diff --git "a/designv11-80.json" "b/designv11-80.json" new file mode 100644--- /dev/null +++ "b/designv11-80.json" @@ -0,0 +1,11154 @@ +[ + { + "image_filename": "designv11_80_0003770_case48305.2020.9216780-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003770_case48305.2020.9216780-Figure1-1.png", + "caption": "Figure 1. Schematic structure of a hydraulic SEA.", + "texts": [ + " The P-ATS compensator enables a more accurate and faster response for hydraulic SEA owing to the reduction in phase drift and amplitude error. The remainder of this paper is organized as follows. The admittance control scheme and problem description are presented in Section 2. The proposed P-ATS compensator is developed in Section 3. The simulation experiments are employed to verify the proposals in Section 4. Finally, the conclusion and future work are provided in Section 5. The definitions of all parameters are given in Table 1. The schematic structure of hydraulic SEA is presented in Fig.1. The load position can be controlled by servo valve via input current I. Thus, the load-side dynamic from cylinder extension Da to load position Dl can be written as: 2- ) 0r a l s L lF D D K M D s \uff08 (1) where Fr represents the external force. ML and Ks represent the load mass and spring stiffness, respectively. Thus, the load nominal dynamic model Pln(s) is derived as Pln(s) = /l aD D . The authors have previously proposed a high-precision position control method that is made independent of hydraulic dynamics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003647_j.procir.2020.09.027-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003647_j.procir.2020.09.027-Figure2-1.png", + "caption": "Fig. 2. Experimental setup with the laser scanner optic and the clamping device. The parts were aligned with an overlap of 12 mm and the bar in the middle of the overlap.", + "texts": [ + " This printer has a heated glass build plate and a resolution of 12.5 \u00b5m in the Xand Y-axes and 2.5 \u00b5m in the Z-axis. The utilized nozzle has a diameter of 0.4 mm. The materials for the transparent and the absorbent parts are natural and black polylactide respectively from Filamentworld. Cura, the software from Ultimaker, was used for slicing the CAD models in 200 \u00b5m thick layers. A layer has an outline that consists of three filaments. The layer infill consist of filaments aligned in the 45\u00b0 direction to the outline (cf. Fig. 2). The orientation of the layer infill changes by 90\u00b0 every layer. For the laser welding experiments, transparent and absorbing parts with 50 mm length, 25 mm width, and 2 mm thickness were produced. To reduce cavities inside the weld seam, bars of 2 mm width and 200 \u00b5m or 400 \u00b5m height respectively were J. Kuklik et al. / Procedia CIRP 00 (2020) 000\u2013000 3 added onto the top layer of the absorbing parts in order to provide additional material in the welding zone (cf. Fig. 2). The printing direction of the bar was parallel to the weld seam. For the laser transmission welding, Laserline\u2019s diode laser LDM300-40 with a maximum power of Pmax = 300 W and a wavelength of \u03bb = 940 nm was used. The laser beam was guided by an optical fiber to a scanner equipped with a focusing optic. This optic generated a focal diameter of 2 mm. The parts were aligned with an overlap of 12 mm, see Fig. 2 right. To generate the required clamping pressure, a pneumatic cylinder pressed the parts against a glass plate with a pressure of 0.2 MPa, see Fig. 2 left. With the parameters for the contour welding process recommended in an earlier study [11], the bars overheat. Therefore, in this investigation the parts were fused by contour welding with scanning velocities ranging from v = 5 mm/s to v = 20.0 mm/s and a laser power range of P = 10.0 W to P = 20.0 W. 3. Results and discussion The absorbent parts with the intermediate layers in form of bars on top were fused in overlap to transparent parts. To study the influence of the different amount of additional material on cavities inside the weld seam, welds were performed with different scanning velocities and laser powers", + " A weld seam with constant width and without cavities is required for a high-quality laser transmission welding of additive manufactured parts. Nomenclature El energy per unit length LA laser absorbent part LT laser transparent part \u03bb wavelength P laser power v scanning velocity 2. Experimental Set-up In this study of laser transmission welding of additive manufactured components, parts were manufactured with a 3rd generation Ultimaker FDM desktop printer. This printer has a heated glass build plate and a resolution of 12.5 \u00b5m in the Xand Y-axes and 2.5 \u00b5m in the Z-axis. The utilized nozzle has a diameter of 0.4 mm. Fig. 2. Experimental setup with the laser scanner optic and the clamping device. The parts were aligned with an overlap of 12 mm and the bar in the middle of the overlap. The materials for the transparent and the absorbent parts are natural and black polylactide respectively from Filamentworld. Cura, the software from Ultimaker, was used for slicing the CAD models in 200 \u00b5m thick layers. A layer has an outline that consists of three filaments. The layer infill consist of filaments aligned in the 45\u00b0 direction to the outline (cf. Fig. 2). The orientation of the layer infill changes by 90\u00b0 every layer. For the laser welding experiments, transparent and absorbing parts with 50 mm length, 25 mm width, and 2 mm thickness were produced. To reduce cavities inside the weld seam, bars of 2 mm width and 200 \u00b5m or 400 \u00b5m height respectively were J. Kuklik et al. / Procedia CIRP 00 (2020) 000\u2013000 3 added onto the top layer of the absorbing parts in order to provide additional material in the welding zone (cf. Fig. 2). The printing direction of the bar was parallel to the weld seam. For the laser transmission welding, Laserline\u2019s diode laser LDM300-40 with a maximum power of Pmax = 300 W and a wavelength of \u03bb = 940 nm was used. The laser beam was guided by an optical fiber to a scanner equipped with a focusing optic. This optic generated a focal diameter of 2 mm. The parts were aligned with an overlap of 12 mm, see Fig. 2 right. To generate the required clamping pressure, a pneumatic cylinder pressed the parts against a glass plate with a pressure of 0.2 MPa, see Fig. 2 left. With the parameters for the contour welding process recommended in an earlier study [11], the bars overheat. Therefore, in this investigation the parts were fused by contour welding with scanning velocities ranging from v = 5 mm/s to v = 20.0 mm/s and a laser power range of P = 10.0 W to P = 20.0 W. 3. Results and discussion The absorbent parts with the intermediate layers in form of bars on top were fused in overlap to transparent parts. To study the influence of the different amount of additional material on cavities inside the weld seam, welds were performed with different scanning velocities and laser powers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003824_aiea51086.2020.00073-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003824_aiea51086.2020.00073-Figure5-1.png", + "caption": "FIG. 5 Load branch cable and state to be installed", + "texts": [], + "surrounding_texts": [ + "The working process of the branch cable installation and clamping device mainly involves three steps: the peeling of the insulating rubber shell of the main cable, the installation of the J-type clamp and the branch cable, and the connection of the branch cable and the main cable. Since the J-type clamp needs to be installed on the bare cable, the insulation layer on the cable is first removed by an automatic stripper, and then the pre-installation of the branch cable and the J-type clamp is performed, as shown in Figures 5 and 6;First install the branch cable, adjust the rotating handle to move the moving crossbar to the proper position, then pull down moving crossbar[9-10]. At this time, the branch cable clamping arm swings accordingly. When the turning arm swings around the hinge point, the connector on the turning arm pulls the spring, and the tension spring simultaneously pulls the branch cable clamping arm, and the branch cable clamping arm tightly bears the branch cable. With the turning arm continues to rotate, the contact point between the connector and the extension spring will pass the dead point formed by the turning arm, the extension spring, and the branch cable clamping arm, thereby ensuring The turning arm will not be pulled back to the starting point. After crossing the dead point, the turning arm leans steadily against the stop crossbar. As shown in Figure 6, the entire locking process is completed, and the branch cable is tightly locked in the wire groove of the clamp. So as to ensure the smooth completion of the entire branch cable installation. Finally, move the entire assembly to the stripping position of the main cable through a long insulating rod, start the motor to connect the main cable and the branch cable, and finally pull the separation wire to realize the separation of the clamping device and the J-type clamp, and the installation work is completed. IV. ADVANTAGES OF BRANCH CABLE INSTALLATION CLAMPING DEVICE The main advantages include the following points; (1) Compared with the existing high-voltage branch cable clamping device, the device has the characteristics of strong 323 Authorized licensed use limited to: University of Canberra. Downloaded on November 12,2020 at 20:55:03 UTC from IEEE Xplore. Restrictions apply. application performance, simple and practical device, light weight and so on. (2) The device is installed by the operator through the insulating rod to avoid the close-range high-voltage electrical work, which increases safety. At the same time, the separation process of the clamping device and the installed J-type clamp is also optimized. The operation is simple and stable. (3) The actual use for many times proves that the clamping force of the J-type clamp after installation of the device is stable and the operation is fast, which greatly improves the work efficiency. V. CONCLUSION Aiming at problems of the current branch cable installation and clamping device, a clamping device based on J-type clamp was designed and developed. This solution uses a non-fixed position branch cable clamping method, which is very suitable for the installation and use of J-type clamp branch cables, and has good adaptability to J components and wires of different wire diameters. The actual field application proves that the device has adjustable clamping force, simple operation and high practical value." + ] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure4.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure4.6-1.png", + "caption": "Figure 4.6 Diagram of the vibration stand.", + "texts": [ + "11) \u0394 = \ud835\udefe0 h g {\ud835\udc4e\ud835\udc4f4 Y 2 1 (b) \u2212 \u2211 i,j=1,3 (xi+1 \u2212 xi) [ y 4 (Y 2 1 (y) \u2212 2 \ud835\udefd4 1 Y \u2032 1(y)Y \u2032\u2032\u2032 1 (y) + Y \u2032\u2032 1 2(y)M M ] yj+1 yj } + S\u2211 k=1 MkY 2 1 (xkyk). For the purpose of verifying the accuracy of the obtained formula (4.2.11), experimental studies were carried out using the resonance method, whereby the fundamental frequencies of cantilever plates with rectangular openings were determined. Studies were conducted on the vibration stand of the excited harmonic oscillations in the range of frequencies from 0.1 to 40Hz (Figure 4.6). The block diagram of connection of the measuring equipment is given in Figure 4.7. Four square plates were subjected to the test: continuous and with 1, 4, and 16 square openings (Figure 4.8). The sizes of the plates are: a = b = 200 mm, h = 1 mm. The material of the plates are steel 260 4 O S C I L L A T I O N O F P L A T E S A N D S H E L L S 4.2 Experimental and Theoretical Research of Oscillation 261 with: E = 2.1\u00d7 106 k\u0393/cm2, and \ud835\udefe0 = 7.8\u00d7 10 k\u0393/cm3. The values of the concentrated masses are identical and equal: M = Q g = 980\u22121 \u00d7 10\u22122k\u0393f 2\u2215cm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001806_j.mechmachtheory.2019.103735-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001806_j.mechmachtheory.2019.103735-Figure7-1.png", + "caption": "Fig. 7. a. Scheme of 3-platform mechanism. n i is the number of legs connecting the two specified platforms. b. A 3-platform gripper. c. A platform that can be represented as a 3-platform mechanism.", + "texts": [ + " If this platform has 6 leg lines, its singularity can be determined by geometrical means. If some of its legs are screws, the determination of singularity will be made by algebraic or combined geometrical and algebraic means. The building of the reciprocal platform is performed by building lines or screws, each of which is reciprocal to each leg. A class of complex mechanisms could be modeled as three platforms connected by legs to one another and to the ground. We denote the ground as 0 and the platforms as I, II and III . The scheme of this family of mechanisms is shown in Fig. 7 a. Two examples of mechanisms that could be modeled as 3-platform mechanisms appear in Fig. 7 b and 7 c. The robot in Fig. 7 b can be used as a gripper. It has actuated legs, while the platforms I, II and III are connected by universal joints, each of which is equivalent to 4 \u201clegs\u201d. The platform in Fig. 7 c has two actuator legs and two actuators with universal joints, which in turn are connected to sliding actuators on the base. Each universal joint is equivalent to 4 \u201clegs\u201d. As seen in these examples, the \u201clegs\u201d are not given explicitly but deduced from the nature of the joint. When checking the singularity, all the actuators are locked. In the examples in Fig. 7 b and c, the bodies referred to as the platforms are marked as I, II and III . The mechanism presented in Fig. 7 c shows that a complex parallel mechanism can be viewed as multi-platform. It is a single platform mechanism with complex kinematic chains connecting the platform I through the legs II and III with the ground. Since the mechanism consists of three platforms which have together 18 DOF, it needs 18 legs, or in general, constraints. The number of legs in connection i is denoted by n i . In the singular position we have forces in all the legs as well as ISAs in each body: k \u2211 i =1 n i \u2211 j=1 f j l j = 0 (10) where the summation is performed over all the legs connected to each platform, and k is the number of platforms adjacent to the given one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001726_icems.2019.8921944-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001726_icems.2019.8921944-Figure4-1.png", + "caption": "Fig. 4. The simplified model of the motor", + "texts": [ + " Demagnetization potential\uff1a When the stator magnetic potential coincides with the direct axis of the rotor, the demagnetization potential is the largest; when the stator magnetic potential is offset from the rotor magnetic potential, the stator magnetic potential decreases along the demagnetization direction component, and the demagnetization is weakened when the stator magnetic potential is intersected with the rotor. When the axial magnetic circuit is facing, the demagnetization potential of the magnetic field of the stator is zero. When the stator demagnetization potential is directly opposite to the position of the magnetic steel, the simplified model of the motor is shown in Fig. 4. For the demagnetization magnetic circuit model at the time of offset, the paper will focus on the analysis. In the model of Fig. 4, the outer magnet of the rotor is equivalent to a permanent magnet magnetic potential source H2l3, and the inner layer is larger. The magnetic steel is equivalent to three magnetic potential sources H1l2, H1l1 H1l2, wherein the magnetic potentials on both sides are equal; H2l3 is connected in series with H1l1, and then connected in parallel with H1l2 on both sides; the equivalent magnetic circuit model is shown in Fig. 5. Rg1, Rg2, Rg3-the figure-air gap reluctance at different positions; Rm1, Rm2, Rm3\u2014magnet resistance at different positions; H2l3, H1l2, H1l1, H1l2-permanent magnet working magnetic potential; F-external demagnetization potential ; \u03c6m1, \u03c6m2, \u03c6m3-the magnetic flux flowing through the permanent magnet" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002365_sii46433.2020.9026258-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002365_sii46433.2020.9026258-Figure2-1.png", + "caption": "Fig. 2. The kinematic model of the dual-arm mobile manipulator: (a) the 5-DOF manipulator equipped with the gripper, (b) the coordinate of the robot base", + "texts": [ + ", \u03b810) T : joint angles of the manipulator equipped with the RGB-D camera \u03b8\u0307a1 = (\u03b8\u03071, \u03b8\u03072, ..., \u03b8\u03075) T : joint velocities of the manipulator equipped with the gripper \u03b8\u0307a2 = (\u03b8\u03076, \u03b8\u03077, ..., \u02d9\u03b810) T : joint velocities of the manipulator equipped with the RGB-D camera px \u2032 b , py \u2032 b : the position of the robot base \u03b8b : the orientation of the robot base vx \u2032 b , vy \u2032 b : the velocity of the robot base \u03b8\u0307b : the rotational angular velocity of the robot base \u03b86, \u03b87, ..., \u03b810 are located from the base to the edge of the manipulator likewise \u03b81, \u03b82, ..., \u03b85 as shown in Fig.2 (a). In this section, cost functions corresponding to each task are defined. Then, these functions should be differentiable. 1) Approaching the target As shown in Fig.3, the cost function of approaching the target, Harm(\u03b8) is given by larm(\u03b8) which is the distance between the base of the manipulator equipped with the gripper and the target as follows, Harm(\u03b8) = |larm(\u03b8)\u2212 carm|, (6) where carm is constant such that 0 < carm. Harm(\u03b8) will take becomes smaller as larm(\u03b8) becomes closer to carm. 177 Authorized licensed use limited to: Murdoch University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-Figure9-1.png", + "caption": "Fig. 9. Design of bending mechanism (a) Arrangement of motors, and (b) Rope path", + "texts": [ + " We previously performed a bending motion experiment with a predetermined rope tension and showed it is possible to bend the elastic telescopic arm [11]. However, comparison of controllability by rope tension and by the rope winding length was not examined. Also, there was no quantitative discussion about the tip position accuracy of the arm. Therefore, the tip position of the arm is controlled by the rope winding length or rope tension, and the controllability is compared in this section. We also control the tip position of the arm using the obtained results and discuss the accuracy quantitatively. We developed the bending mechanism [11]. Fig. 9a shows the arrangement of motors in tendon-driven part. The coordinate system is defined as a right-handed system with the longitudinal direction of the arm as the z direction and the vertically upward direction as the y direction. In this driving method, the arm can be bent on the x-y plane by the combined force of three rope tensions. In addition, among the methods that can move on the x-y plane and always generate a force in the y direction, this method can minimize the number of driving parts", + " There are studies of bending motion by using ropes [12][13]. In these studies, ropes are placed along the arm using rope guides, and the bending moment is generated by winding up the rope attached the tip position of arm. However, when using a telescopic structure in which the rigidity of the arm is high and the diameter increases by the root compared with these studies, there is a problem that only the tip is bent by winding the rope attached to the tip. So we created a rope path as shown in Fig. 9b. This is a method of reciprocating the rope by pulleys. The rope tension can be applied at multiple positions of the structure since the rope tension is applied to all passing pulleys. For this reason, it is possible to obtain a distribution of bending moment that increases continuously toward the root by changing the arrangement of the pulleys in the axial and radial directions of the structure. In addition, it is possible to wind all the ropes at the time of contraction. First, in order to compare the controllability of the tip position with the rope winding length and the rope tension, we performed the bending experiment in the y direction of the arm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001879_012007-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001879_012007-Figure4-1.png", + "caption": "Figure 4. Illustrate of grasping candidates obtained using our algorithm", + "texts": [ + " In the pretreatment step, we have simplified the main part of target object. We set the main part as the region of interest and constrained the grasping pose based on the simplified shape features to obtain the grasp candidates. The following experiment confirmed this method is effective for robot grasping. The grasp constraint is shown below: Simplified into sphere: the center position cenp of the gripper is located in the center of the ball, the orientation cenr of the gripper is a random direction vector (shown in Figure 4(a)). Simplified into ellipsoid: the center position cenp of the gripper is located in the central long axis of the ellipsoid, the orientation cenr is the same as one of the two short central axis (shown in Figure 4(b)). Simplified into cylinder: the center position cenp of the gripper is at the central axis, the orientation cenr is parallel to the central axis (shown in Figure 4(c)). Simplified into parallelepiped: the center position cenp of the gripper is at the center long axis, the orientation cenr is the same as one of all edges except the shortest edge of the parallelepiped (shown in Figure 4(d)). We define D(h) as the volume occupied by the robotic gripper when it is fully open. Defined closing region ,C(h) , as the volume area swept by the gripper when the fingers are closed. In the vertical direction of each gripper, we \u201cpush\u201d the gripper until the following conditions are met, and Figure 3 shows the above conditions occurred when grasping object. Condition 1. The grabs hand and the point cloud will not collide and contact during the process of \u201cpush\u201d hand : D(h) C , and the center of the C(h) is cenp ", + " (a) Condition 1 occurred when grasping. (b) Condition 2 occurred when grasping. (a) (b) Figure 3. Illustrate of the conditions present in the paper. (a) Condition 1 occurred when grasping (b) Condition 2 occurred when grasping On the basis of grasping constraints, considering the noise of the input cloud and obtaining sufficient grasp candidates, all the gripper positions cenp have random errors of 0~1 cm and all the gripper orientations cenr have random errors of 0~15 degree. As illustrated in Figure 4, we obtain grasping candidates though the above grasp constraints. ICMME 2019 IOP Conf. Series: Materials Science and Engineering 717 (2020) 012007 IOP Publishing doi:10.1088/1757-899X/717/1/012007 The grasp training network in this paper is defined as a binary classification task that the input is the feature representation of a set of grasp candidates and the output is a prediction of whether or not the grasp candidate is graspable. The network structure we use is the same as the LeNet-5 [23] CNN structure: three convolutional layers, two pooling layers and a fully connected layer with a softmax on the output" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.9-1.png", + "caption": "FIGURE 7.9", + "texts": [ + " As f is always less than 1, the surface tends to be hydrophobic with the increase of surface roughness. For homogeneous hydrophobic surfaces, Eq. (7.11) is often used to explain why solid surfaces are hydrophobic. Based on the Wenzel and the Cassie Baxter models, the relationships between the geometrical parameters and the contact angle of the anisotropic microgroove textured surface and the square microconvex surface were solved, respectively, and the influence of the geometrical parameters of the surface on its wettability was discussed. Fig. 7.9 shows the topography of the isotropic microconvex surface. The microconvex width is set to a (\u03bcm) along both horizontal and vertical directions; the width of the groove is set to b (\u03bcm) in both horizontal and vertical directions; and the microconvex width is set to h (\u03bcm). For the Wenzel model, the roughness factor r is available. r5 4ah a1b\u00f0 \u00de2 1 1 (7.12) For the roughness factor f of The Cassie Baxter model, its expression is as follows: f 5 a2 a1b\u00f0 \u00de2 (7.13) Topography of isotropic microconvex surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001924_icoecs46375.2019.8949914-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001924_icoecs46375.2019.8949914-Figure2-1.png", + "caption": "Fig. 2. EM topology with a tooth-coil winding according to the alternate-teeth-wound scheme", + "texts": [ + " All these methods are based on the redundancy of the EM elements, which can lead to an increase in the mass or to an increased thermal load on the windings, in the event of failure of one of the elements. It should be noted that the main requirement for implementing a fault-tolerant EM (multi three-phase system) is thermal, galvanic and electromagnetic isolation of the EM windings. This requirement can be realized only by applying the toothcoil winding according to the alternate-teeth-wound scheme, Fig. 2. To select an effective way to ensure the fault-tolerance of the EM, it is advisable to consider all the above mentioned features depending on the design topology and make a detailed analysis of multi-phase and duplex three phase systems to choose the most effective and efficient option. In general, these methods can be categorized in integer and fractional multiplicity. The first is considering that the failure of the X element is possible only when X + 1 of the same type has failed. In this case, the reliability of the EM is defined as [38]: ( )( ) 1 1 ( ) 0 i x p t p ty i i = = \u2212 \u2212\u220f = , (1) where ( )yp t is a probability of no-failure operation of entire EM; ( )ip t is a probability of no-failure operation of an EM element; \u0445 is either the number of reserved elements or number of duplicated elements in EM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001903_j.jmmm.2019.166372-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001903_j.jmmm.2019.166372-Figure11-1.png", + "caption": "Fig. 11. The schematic picture of a ring core with four interlocks. The interlocks are parallel to flux.", + "texts": [ + " (g) The front of a laminated core with interlocks parallel to the long- itudinal direction. (h) The back of a laminated core with interlocks parallel to the long- itudinal direction. In this chapter, the modelling method for a magnetic circuit (a ring core in this case) with interlocking [6] is briefly explained and the magnetic characteristic of the damaged region for this model is calculated. First, it is assumed that the interlocked ring core is represented by a magnetic series circuit with undamaged and damaged regions as shown in Fig. 11. If there is no leakage of flux from the ring core, the measured permeability and iron loss are expressed by the following equations. = + \u2212l \u03bc Nd \u03bc l Nd \u03bc'' 'a (6) = + \u2212W Nd l W l Nd l W 'a '' (7) where \u03bca and Wa are the measured permeability and iron loss. \u03bc'' and \u03bc' are the permeability of damaged and undamaged regions. W\u2019\u2019 and W\u2019 are the iron losses of damaged and undamaged regions. N is the number of interlocks in a ring core. l is the average length of the entire magnetic circuit, which is approximated by the following equation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001739_ismsit.2019.8932756-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001739_ismsit.2019.8932756-Figure2-1.png", + "caption": "Fig. 2. Lead Screw and Stepper Motor", + "texts": [], + "surrounding_texts": [ + "978-1-7281-3789-6/19/$31.00 \u00a92019 IEEE\nKeywords\u2014 automation, gardening, mobile application, timely monitoring\nI. INTRODUCTION\nIn manufacturing sector, to control machine tool a processed is used which is called CNC machining. The process includes the use of computers to control the machine tools. Tools can be any object which helps us in performing different task like extruder in 3D printer, blade in lathe machine, drill bit in milling machine. CNC stands for Computer Numerical Control. It look like an ordinary personal computer (PC), but it differs due to unique software and console to control the machine.\nTo control the speed, location, coordinates and feed rate of machine a specialized CNC machining language is used, which is known as G-code. With the help of CNC machining we can control the position and velocity with great precision. CNC machining is currently being used in manufacturing for plastic and metal parts.\nThere are lots of advantages of using CNC machines. It\u2019s hard to manufacture any part in manual machining. Manual machining consumes more time and energy and it does not have much precision. While using CNC machining a lot of man power and time is secured and also the parts manufactured with CNC machining has great precision.\nA CNC machine consist multiple individual motors, which helps its tool to move in respective specific direction. A CNC Machine consist of two or more axis of motion. Usually CNC machine which are being used are consist 3-axis of motion. If there is any additional rotational movement then it will be 4-axis.\nII. LITERATURE REVIEW\nCommercially, Agriculture has been reached a newly high level of automation up till now, mainly for growing\ncrops on broad land. Fine-grain satellite imagery is also available commercially for pesticides and fertilizer related applications, leading to the precision agriculture\u2019s novel paradigm, its basic goal is to save water and pesticides [1]. The interest is growing day by day in autonomous farming so we can move towards smaller robotic platforms that can work on individual basis and can done some manipulation in the field with the help of precise sensing [2].\nWith the increasing research in the field of computer vision [3] and mechatronics which is helping us to led to an autonomous solution for harvesting some specialty crops which includes Cherries [4], apples, tomatoes, cucumber, mushrooms, strawberries, melons and much other.\nAnother active research on which work is being done is automatic weed control. Grey-level vision is used to navigate in structured outdoor environment and uses color vision to differentiate between the required crops and weed which is needed to be removed. Beside their multiple applications in the field, we envision automated agriculture robot which is precise and able to work without any operator in any sort of environment like urban areas, house roofs or can easily work in harsh environment like outer space.\nAlso, a lot of work is done in this field like autonomous targeted spraying [5] to removes pests from crops which can destroy and can have an effect on crops production, design of optimum manipulator for autonomous de-leafing [6] process of cucumber plant to prevent it from fungal disease which is considered by its growers but is costly. In addition to all of this there are also some virtual experimentation frameworks which have been developed for agriculture robots, [7], [8] ForboMind is a customized software platform which was introduced to support and help field robot done agriculture task done with precision and to promote to reuse the robotics components.\nOn the other hand, Agriculture field robots contributes to improve soil health, yield and reliability of operation. Which are commonly equipped with multiple sensors and cameras for navigation, localization, mapping and path planning algorithm.\nAlso, farming industry make use of drones for surveillance and monitoring fields. These drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides.", + "III. SOCIO-ECONOMIC SIGNIFICANCE\nAt present farming industry make use of drones for surveillance and monitoring fields these drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides. Similar Machinery exists but is only household specific; however we are aiming to target research and development departments in pharmaceutical and food industry, where setting up small greenhouse is required. Further we would be using some tool head for multiple operations instead of relying on magnetism for tool selection. Cutting down cost is another objective by using alternative materials to steel.\nIV. DESIGN\nWhile developing the project we went through some major and minor issues which lead us to make multiple iterations in the design phase. In this section we will discuss these iterations.\n Iteration 1:\nThis is the initial idea and concept on which we started to work. Our initial idea was to have x-axis, y-axis, z-axis and one rotary axis for the axis of motion. We also thought about adding multi-head mounter and add moisture sensor, shower and seeder. Initially we want to use rack and pinion mechanism. Issues: Rack and pinion mechanism can increase the vibrations which can affect the structure.\nWe then start to work on to select a mechanism for x-axis. We start to work on the stepper motor and lead screw mechanism. Issues: Our project has an open base so lead screw can\u2019t be mounted at the bottom like every CNC milling machine and usage of two stepper motor for just one axis can effect or destroy the structure if something happens to one motor.\n Iteration 3: So we started to think more about it and started to work on another solution to move the x-axis with just one motor. We finally decide to move the whole axis with one motor by using belt and pulley mechanism as shown in figure below.\nIn this project we use t-slotted extruded bars for developing our structure. Y-axis consist of simple lead screw mechanism.\nZ-axis consist of simple lead screw mechanism for linear motion. On z-axis a servo motor is mounted for giving a rotary motion to the multi head. The final design after the iterations is shown below.", + " Vacuum Pump: When a signal is sent out from controller to the vacuum pump and vacuum pump turn on and suck air and move towards seed when it come near the seed, then seed float and got stuck on the nozzle until the vacuum is on. When it reaches to the desired location then the controller sends another signal and vacuum pump turns off and the seed drops on the location the grid.\n Water Pump: When a signal is sent out from the controller to the water pump then water pump turn on and sprinkle the water over the grid with the help of shower for some specific time and then another signal comes from the controller and turns it off. Piping and instrumentation diagram (P&ID) is shown below:\nCurrently there are lots of global challenges which we are facing like global warming, food needs, poverty, degradation of climate and much more. Sustainable development goals introduced by United Nation helps us address these problems and base on these problems find solution to make future better not just for humans but also for the creature that exist on globe. The goals which can be achieved by our project \u201cGarden Tech\u201d are as follows:\n Zero Hunger: As we all know that human population is increasing rapidly to accommodate them deforestation is taking place. It is effecting our climate and results in global warming due to which glaciers are melting. Flooding is taking place more often than ever before and leads to devastation of arable land. Due to these circumstances food growth is not increasing rapidly like human population. The anticipated population according to United Nations Department of Economic and Social Affairs (UN DESA) is shown below:" + ] + }, + { + "image_filename": "designv11_80_0003434_tcsme-2019-0259-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003434_tcsme-2019-0259-Figure2-1.png", + "caption": "Fig. 2. A finite element model of the housing. Image source Ren et al. 2018; use permitted under the Creative Commons Attribution license CC BY 4.0. [Colour online.]", + "texts": [ + " The two-node form stiffness matrix yields KBr = Km \u2212Km \u2212Km Km \u00f05\u00de where Km is the single node form of the stiffness matrix with six DOF. Given the damping characteristics and external force, the impedance equation of the shaft subsystem can be expressed as follows: ZBr\u00f0\u03c9\u00deVBr\u00f0\u03c9\u00de = FBr\u00f0\u03c9\u00de\u00f06\u00de where ZBr (\u03c9)=CBr+KBr / (j\u03c9) is the impedance matrix of bearing subsystem, CBr is the bearing damping; VBr is the velocity vector of bearing; and FBr is the external force vector of bearing. A finite element (FE) model is built in ANSYS software, as is shown in Fig. 2. The material is steel, the density is 7850 kg/m3, the Young\u2019s modulus is 2.07 \u00d7 1011 Pa, Poisson\u2019s ratio is 0.3, and the damping ratio is 2%. The four-node tetrahedron element is used, and the element size is 5 mm. The FEmodel consists of about 76 000 nodes and 315 000 elements. Four bearing holes are coupled to the central nodes through distributed force coupling. Four bearing holes central nodes and six isolator connection nodes constitute 10 external nodes. The block Lanczos method is used to process modal analysis, and the first 500 modes are extracted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002215_s00170-019-04806-8-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002215_s00170-019-04806-8-Figure3-1.png", + "caption": "Fig. 3 Projections of the velocities of the planar motion onto axis systems \u041e1\u0443, \u041e1z, and \u041e2\u03b7, \u041e2\u03b6", + "texts": [ + " NKx \u00bc NKcos\u03b3cos\u03c8; NKy \u00bc NKcos\u03b3sin\u03c8; NKz \u00bc NKsin\u03b3: \u00f08\u00de Lines of action of normal reactions NB1, NB2 at symmetric points of contact \u04121 and \u04122 pass through point \u041e of peg and hole axes\u2019 intersection [18], and their direction cosines shall be defined from geometrical ratios. They should have the following values after transformations: cos\u03b1N B1 \u00bc \u2212 S1cos\u03c8\u00fe bsin\u03c8 B1 ; cos\u03b1N B2 \u00bc \u2212 S1cos\u03c8\u2212bsin\u03c8 B1 ; cos\u03b2N B1 \u00bc \u2212 S1sin\u03c8\u2212bcos\u03c8 B1 ; cos\u03b2N B2 \u00bc \u2212 S1sin\u03c8\u00fe bcos\u03c8 B1 ; cos\u03bbN B1 \u00bc a1 B1 ; cos\u03b3NB2 \u00bc a1 B1 ; \u00f09\u00de where (Fig. 3) b \u00bc BB1 \u00bc BB2; B1 \u00bc OB1 \u00bc OB2 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 0:25D2 \u00fe a21 q : Hence, projections of normal reactions shall be written as follows. Nx B1 \u00bc NB1cos\u03b1 N B1; Nx B2 \u00bc NB2cos\u03b1 N B2; Nx K \u00bc NKcos\u03b1 N K ; Ny B1 \u00bc NB1cos\u03b2 N B1; Ny B2 \u00bc NB2cos\u03b2 N B2; Ny K \u00bc NKcos\u03b2 N K ; Nz B1 \u00bc NB1cos\u03bb N B1; Nz B2 \u00bc NB2cos\u03bb N B2; Nz K \u00bc NKcos\u03bb N K : The direction of friction force is always opposite to the absolute velocity of the point of its application that involves determining the values of velocities of contact points of parts \u04121, \u04122, and \u041a", + " The velocities of the points of the peg located in the plane symmetry are defined as rotational around the instantaneous center of velocities, located at the intersection of the perpendiculars to the velocities VK and VA. Velocities of symmetric points of contact \u04121 and \u04122 are equal to point \u0412 velocity, since they are located on one perpendicular line to the symmetry plane, passing through point \u0412, hence, the values of the contact points\u2019 velocities shall be equal to V\u03b3 B1 \u00bc V\u03b3 B2 \u00bc V\u03b3 B \u00bc BL \u03b3 ; V\u03b3 K \u00bc KL \u03b3 : Projections of velocities V\u03b3 B1; V \u03b3 B2;V \u03b3 K of points \u04121, \u04122, and \u041a on movable axes \u041e2\u03b7 and \u041e2\u03b6 are equal to (see Fig. 3) V\u03b3 B1\u03b7 \u00bc V\u03b3 B2h \u00bc \u2212BL \u03b3 cos\u03b52 \u00bc \u22122a2 \u03b3 ; V\u03b3 B1\u03b6 \u00bc V\u03b3 B2\u03b6 \u00bc \u2212BL \u03b3 sin\u03b52 \u00bc \u22122S2 \u03b3 : V\u03b3 K\u03b6 \u00bc KL \u03b3 ;V\u03b3 K\u03b7 \u00bc 0: \u00f010\u00de Projections of these velocities V\u03b3 B1;V \u03b3 B2;V \u03b3 K to auxiliary axis \u041e1h are equal (see Fig.4) to V\u03b3 B1h \u00bc V\u03b3 B2h \u00bc \u2212BL \u03b3 cos\u03b51 \u00bc \u22122a1 \u03b3 ; V\u03b3 Kh \u00bc \u2212KL \u03b3 sin\u03b3: \u00f011\u00de These expressions include (see Fig.3) \u03b51\u2014angle between segment \u0412L and axis of hole, \u03b52\u2014angle between segment \u0412L and peg axis, cos\u03b51 \u00bc a1 OB; sin\u03b51 \u00bc S1 OB; sin\u03b52 \u00bc S2 OB; cos\u03b52 \u00bc a2 OB; B \u00bc OB \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a21 \u00fe S21 q \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a22 \u00fe S22 q ; BL = 2OB = 2B. Projections of velocities V\u03b3 B1; V\u03b3 B2; V \u03b3 K to fixed axes\u041e1\u0445, \u041e1\u0443, and \u041e1z (Fig. 4) are equal to V\u03b3 B1x \u00bc V\u03b3 B2x \u00bc \u2212V\u03b3 B1hcos\u03c8 \u00bc \u22122a1cos\u03c8\u03b3 ; V\u03b3 K\u0445 \u00bc \u2212KLsin\u03b3cos\u03c8\u03b3 ; V\u03b3 B1y \u00bc V\u03b3 B2y \u00bc \u2212V\u03b3 B1hsin\u03c8 \u00bc \u22122a1sin\u03c8\u03b3 ; V\u03b3 Ky \u00bc \u2212KLsin\u03b3sin\u03c8\u03b3 ; V\u03b3 B1z \u00bc V\u03b3 B2z \u00bc \u22122S1 \u03b3 : V\u03b3 Kz \u00bc KL \u03b3 cos\u03b3: \u00f012\u00de Rotation about hole axis occurs with angular velocity of \u03c8\u0307 \u00bc d\u03c8 dt " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002128_amm.896.183-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002128_amm.896.183-Figure2-1.png", + "caption": "Fig. 2. Multi-function sensing tool. (a). The depiction of the sclera force and the insertion depth, the sclera force is the contact force between the sclerotomy port and the tool shaft, the insertion depth is the distance between the tool tip and the sclerotomy port when tool is inside of the eyeball. (b). the tool dimensions", + "texts": [ + " Other standard sequential models, such as Kalman filters or fixed-lag smoothers and predictors, are ill-equipped to learn long-range dependencies [16]. In this paper, we report initial results for predicting possibly unsafe user behavior immediately before it is about to occur, to provide enough reaction time for preventing tissue damage during robot-assisted retinal surgery. An LSTM model is adopted and trained to predict the user behavior, which is defined through two correlated time-series: the sclera force and the insertion depth as shown in Fig. 2 (a). The predicted information is compared with known safe operation thresholds to give auditory feedback, and to alarm users about any upcoming misoperations. Experiments are conducted on a dry eye phantom to examine the feasibility of the proposed method. The goal of this work is to establish the user\u2019s behavior library, and achieve a safety-aware cooperative control in robot-assisted surgical systems. While there are a number of other signals useful for behavior prediction, such as tool inclination and eye tracking and imaging, this initial investigation is limited to the two basic signals: sclera force and insertion depth", + " During operation both the user and the robot hold the tool, the user\u2019s manipulation force is applied on the robot handle and fed as an input into the admittance control law as shown in Eq. (2) where x\u0307 hh and x\u0307 rh are the desired robot handle velocities in the handle frame and in the robot frame, respectively, Fhh is the user\u2019s manipulation force input measured in the robot handle frame, \u03b1 is the admittance gain tuned by the robot pedal, Adgrh is the adjoint transformation. The multi-function sensing tool is designed and fabricated as shown in Fig. 2. The tool shaft is made of a stainless steel wire with a diameter of 0.635 mm and machined to contain three longitudinal V-shape grooves. Each groove is filled and glued with one optical fiber containing three FBG sensors (Technica S.A, Beijing, China), i.e., nine FBG sensors are embedded in the tool shaft in total. The tool dimensions are shown in Fig. 2. The multi-function sensing tool can measure the sclera force and the insertion depth. Considering ambient temperature and noise, we use the sensor reading \u2206s to calculate the force defined as follows: where \u2206\u03bbi j is the wavelength shift of the FBG, i =I, II is the FBG segment, j = 1, 2, 3 denotes the FBG sensors on the same segment. The sclera force exerted on the tool shaft generates strain in the FBG sensors, thus it contributes to the sensor reading: \u2206Si = KiMi = KiFsdi, (4) where \u2206Si = [\u2206si1, \u2206si2, \u2206si3]T denotes the sensor reading of FBG sensors in segment i, Fs = [Fx, Fy]T denotes the sclera force applied at sclerotomy port, di denotes the distance from the sclerotomy port to FBG at segment i along the tool shaft, Mi = [Mx, My]T denotes the moment attributed to Fs on FBG sensors of segment i, Ki (i = I, II) is 3x2 constant coefficient matrices, which is obtained through the tool calibration procedures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001683_icems.2019.8921881-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001683_icems.2019.8921881-Figure1-1.png", + "caption": "Fig. 1. Initial analysis model.", + "texts": [ + " However, this method can only restrain the torque ripple when the SRM rotates in the forward direction, which is not suitable for the frequent forward and reverse rotation switching occasions. In this paper, a perfectly symmetrical structure of the stator and rotor pole is proposed to suppress the torque ripple in the positive and negative rotating conditions. The torque ripple and the average torque with variable optimization parameters are calculated by 2-D FEM. Then, the multi-objective optimization is carried out by the PR models of the optimization objectives combined with the Pareto genetic algorithm (PGA). II. ANALYSIS OF INITIAL SRM Fig. 1 shows the topology of a 12/8 SRM, which has been taken as the initial motor to be studied. The SRM consists of stator, rotor, shaft and coils of stator winding. The key design parameters are listed in Table I. The torque and inductance profiles at 4000 rpm are obtained by FEM, as shown in Fig. 2. As can be seen from Fig. 2, for an ideal SRM, the inductance profile should be constant prior to the overlap of stator and rotor poles, so no torque is produced in this region. In fact, the inductance would nonlinearly increase prior to the overlap angle due to the fringing flux effect and local magnetic saturation, as presented in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002640_b978-0-12-821354-4.00011-x-Figure11.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002640_b978-0-12-821354-4.00011-x-Figure11.2-1.png", + "caption": "FIG. 11.2", + "texts": [ + " Their demand is visible through a yearly growth rate of nearly 25% in numerous fields ranging from packing to complicated biomedical applications (Dalton et al., 2002; Thostenson et al., 2005). NCs can be termed solid materials with multiphases, where one or all three phases may have dimensions less than 100nm or Biosensors reported until now for herbicide sensing. structures of the material with multidistance nanoscale repeat spaces (Bordes et al., 2009). The additives in the nanosize and their superior dispersion with the support vary the nanocomposite from the conventional composite (Fig. 11.2). The NC is ecofriendly and has tremendous potential in the aerospace, automotive, electronics, and biotechnology sectors (Choi and Awaji, 2005; Roy et al., 1986). These composites have more phases in the nanometric dimension with large surface areas and enhance interactions at the phase interfaces (Camargo et al., 2009; Schmidt et al., 2002). These exclusive properties are of immense benefit for the effective use of agrochemicals such as pesticides, herbicides, insecticides, fungicides, etc., as nanocomposites" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure10-1.png", + "caption": "Fig. 10. Flux density of proposed model", + "texts": [], + "surrounding_texts": [ + "The proposed model with 6 poles and 48 slots reduces the cogging torque effectively. The significant improvement of performance for the proposed model, as shown in Fig. 11 can be identified by the smallest cogging torque in the beginning 15 tested points in the mechanical rotor degrees (00~150), and the last 15 tested point in the mechanical rotor degrees of (450~600). The maximum cogging torque of the proposed model is 0.007031968 Nm. While the maximum CT at the rotor rotates for the original model is 0.5830463 Nm, then we got the CT reduction as much 98.79%. Compare with the onestep slotting model. It is observed that the CT Reduction as much 93.90% with the maximum CT for this one step slotting is 0.556849257 Nm. It means that the proposed model is highly accepted in this research. At the beginning of rotation from motionless or at low speed, the original model and the model 1 need more mechanical energy to attain the same speed rotation compared with the proposed model (two-steps model). The fluctuation distribution of cogging torque for the proposed model. Authorized licensed use limited to: City, University of London. Downloaded on July 10,2020 at 12:04:30 UTC from IEEE Xplore. Restrictions apply. V. CONCLUSION AND REMARKS The influence of the slot opening width on the pole magnet area was investigated in the paper. From the simulation results, it can be concluded that the smaller crosssection area of magnet pole results in the decrement of cogging torque and air gap normal flux density in the InsetPMMs. Additional slotting in the magnet edge reduces the magnet pole cross-section and the cogging torque. The novelty of the proposed Model achieves to adjust the magnet pole arc and distance without changing the rotor diameter and stator construction." + ] + }, + { + "image_filename": "designv11_80_0001110_3348445.3348476-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001110_3348445.3348476-Figure1-1.png", + "caption": "Figure 1. Structure of a quadrotor.", + "texts": [ + " Finally, the conclusions are presented in Section 6. 2. DYNAMIC MODELING OF QUADROTOR The attitude and position of a quadrotor can be controlled by changing the speeds of the four motors that turn the propellers. The rotation of rotors causes the thrust on the quadrotor. The pitching moment and rolling moment are generated by the difference among four rotors thrust. The yawing moment is caused by the unbalanced of the four rotors rotational speeds. The earth fixed frame and rigid body model of a quadrotor are described in figure 1. As shown in figure 1, two pairs of propellers 1 3( , ) and 2 4( , ) rotate in the opposite direction. Two diagonal motors 1 3( , ) rotate in clockwise direction and the other motors 2 4( , ) rotate in counter-clockwise direction. The roll, pitch, and yaw angles , , can be acquired from the rotation of a quadrotor\u2019s body frame in x, y, and z axis with the control inputs 1U , 2U , 3U , and 4U as follows: 1 2 3 4 4 4 , , , 1 cos cos , sin sin cos sin cos , sin cos cos sin sin . y z r x x x x z r y y y x y z z i i J L U i i i i i J L U i i i i i L U i i z U g m U x m U y m (1) In equation (1), the parameters xi , yi , and zi mean the body inertia in x, y, and z coordinates, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000464_012036-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000464_012036-Figure1-1.png", + "caption": "Figure 1. Transverse vibration model of wire rope.", + "texts": [ + " The model establishment and solution in this paper are based on the following four assumptions: 1. Ignore the influence of airflow and friction in wellbore; 2. Ignore the influence of balanced wire rope vibration; 3. Ignore the influence of torsion and longitudinal vibration of wire rope, and the elastic deformation caused by transverse vibration is much smaller than the length of the entire string. 4. In the operation of high-speed elevator, the linear density of wirp rope is , the cross-sectional area is A ,and the elastic modulus is E remain unchanged. Figure 1 is a transverse vibration model of high-speed traction elevator lifting system. In this paper, the wire rope is regarded as a variable length string with axial motion, the linear density is , the cross-sectional area is A , the elastic modulus is E , and the length l t varies with time t . The lifting weight is simplified as a rigid weight with mass m suspended at the bottom of the chord line, which is free longitudinally, Constrained by a spring with a lateral stiffness of k and a damper with a damping coefficient of c , the existence of a rigid tank passage in a real high-speed elevator hoisting system is simulated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002071_iccas47443.2019.8971752-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002071_iccas47443.2019.8971752-Figure3-1.png", + "caption": "Fig. 3 A 2D-model of the WIP wheelchair with an au tomatic slider.", + "texts": [ + "2 Relationship among acceleration, a tilt angle, and a slider position To know the relationship among acceleratio n, a bod y tilt angle, and a slider position is impor tant to achieve the I (. \u00b7 )2 I . 2D = Dw+ D; = r- Ow - 'l/Jc + 2 CnAn. (7) Then, the Lagrangian function is given by L = 'I'; + Ti, + T\" - Uc. - Us - U\" . Choos ing a generalized co ordination as q = ['l/Jc Ow Aaf and genera lized force as Q = [0 Tw faf , motion equations are derived through the following Lagrangian equations ( I) (2) (4)W [ . ]T0 c \" = 0 'l/Jc 0 . 3.1 Motion equations A planar model of a WIP wheelc hair is shown in Fig. 3 and its model parameters are described in Table I. It con sists of a wheel part, a wheelc hair-body part including a slider, and a seat part including a rider. The wheelchair body part and the seat part are represented as rectangles. The seat is fixed on a top of the slider, where a rider 's CoG locates at a position of (Aa + Am, l,,) with respect to a wheel shaft. Motion equations are derived by La gra nge 's method as follows. The position of CoG , Gb and G\", with respect to the global coo rdinates are given as [ Xc - olbS,pc ] , w P C b = -hC,pc [ Xc + l\"S ,pc + AmaC,pc ] w P c = 0h l\"C,pc - AamS,pc where Aam = Aa + Am and S* = sin (*),C* = cos]\u00bb)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003038_012194-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003038_012194-Figure5-1.png", + "caption": "Figure 5. Simulation diagram of the mountain tractor suspension implements", + "texts": [ + " Therefore, as long as the length of the left telescopic rod (2) and the length of the right telescopic rod (3) are adjusted, the roll angle target \u03b1 and the pitch angle target \u03b2 of the tractor implement attachment frame (5), that is, the roll angle \u03b1 and pitch \u03b2 of the tractor implement is adjustable. An example of a system simulation of a mountain tractor implement is given by the software AME. Figure 3. Analysis diagram of the lower bar degrees of freedom Figure 4. Schematic diagram of the adjustment of the position and posture on the tractor suspension implement ICEMEE 2020 IOP Conf. Series: Earth and Environmental Science 508 (2020) 012194 IOP Publishing doi:10.1088/1755-1315/508/1/012194 Figure 5 shows the adjusting process of the mountain tractor suspension implements on the slope. As the mountain tractor is connected to the rotary suspension implement to perform rotary tillage on the slope, the left telescopic rod can be adjusted by the manual directional valve length. Then the length of the right telescopic rod through the manual reversing valve can be adjusted. It is assumed here that the tilt angle is 15\u00b0, the roll angle of the tractor suspension implement attachment frame is 15\u00b0, that is, the rotary tiller is also in a roll posture of 15\u00b0, satisfying the contouring operation on the slope field" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure33.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure33.1-1.png", + "caption": "Fig. 33.1 a Geometry of the present work, b geometry of the present work after meshing", + "texts": [ + " The Gaussian distributed disk heat source model that is utilized in the present analysis is given by [30]. q \u00bc 3gP pr2 exp 3 x2 \u00fe y2\u00f0 \u00de r2 ; \u00f033:6\u00de where q is the heat flux, \u03b7 is the laser absorption coefficient, P is the laser power, r is the effective laser beam radius, and x and y are the respective position of the laser beam at a given time. A computational model was developed using FE-based ANSYS 14.5 software. The dimensions of the substrate are 10 mm 31 mm 4 mm and five layers with layer thickness 1 mm as shown in Fig. 33.1. Also, temperature-dependent material properties such as thermal conductivity, density and specific heat are used [31]. The scanning pattern of the laser on successive layers is shown in Fig. 33.2. The process variables used in the present work are given in Table 33.1. Six different data sets are utilized to simulate laser additive manufacturing process and to analyze melt-pool dimensions and temperature distribution. The deposition of layers is simulated using the element birth and death technique in ANSYS APDL" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure4-1.png", + "caption": "Fig. 4. Initial/Original Model", + "texts": [], + "surrounding_texts": [ + "Using the FEMM 4.2 and coupled with LUA 4.0 scripting, the PMMs characteristics were investigated [7],[11],[12]-[14]. By the combination of FEMM and LUA 4.0 could increase a quick execution for the implementation of a complete simulation of a specific PMMs. Another advantage of LUA application is the capability of parallel computation can be achieved. At the beginning of each simulation, the simulated of PMMs structure were generated in Auto-CAD then exported to the FEMM file. Comparisons of air gap magnetic flux distribution and CT for the PMMs studied were investigated. This means that the proposed PMM model (twosteps slot) was promising for the presence of two steps of slotting in the magnets that do not distort the balance of magnetic force in the air gap of the machine of one-step model. Figures 4, 5 and 6, shows that the value of the flux distribution due to the changing in the magnet structure will not destroy the value of the machine's core losses, as it could be observed that all the value of the flux density approximately 1.436 Tesla. It has been found that if it implemented to the proposed method, it will not change the core losses of the proposed model." + ] + }, + { + "image_filename": "designv11_80_0003400_sdpc.2019.00093-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003400_sdpc.2019.00093-Figure2-1.png", + "caption": "Fig. 2. Bearings used in the experiment", + "texts": [], + "surrounding_texts": [ + "The wavelet transformation only decomposes the signal into two parts, high frequency and low frequency, which then performs the same operation on the low frequency components. The high frequency components that have been decomposed are not calculated. The low frequency components and the high frequency components are decomposed simultaneously by wavelet packet transformation. Therefore, wavelet packet transform is a more sophisticated analysis method for signals. Define the subspace as the closure space of the function ( ) , as the closure space of function ( ) . Assuming that ( ) \u2208 , can be expressed as = \u2211 , (2 \u2212 ) (1) The wavelet packet decomposition algorithm can be expressed as:use , to find , and , . , = \u2211 , (2) , = \u2211 (3) The wavelet packet transformation can adaptively search for the corresponding frequency band according to the signal characteristics, so that the decomposed detail component matches the spectrum of the signal. The decomposed detail components have the same bandwidth in the spectrum, but the frequency centers are different. So the wavelet packet decomposition can be simply understood as a signal passing through a series of filters with different center frequencies but with the same bandwidth." + ] + }, + { + "image_filename": "designv11_80_0002567_kem.839.73-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002567_kem.839.73-Figure1-1.png", + "caption": "Figure 1 - Block-scheme \u201cVARISKAF-100MBS\u201d: 1 \u2013 PC, 2 \u2013 ytterbium fiber laser LK \u2013 100 \u2013 \u0412, 3 \u2013 collimator, 4 \u2013 scanner, 5 \u2013 heating system, 6 \u2013 vertical movement motor, 7 \u2013 a hopper for powder-coating, 8 \u2013 CNC, 9 \u2013 vacuum chamber, 10 \u2013 working window", + "texts": [ + " (#541000290, University of Wollongong, Wollongong, Australia-18/07/20,23:17:35) In experiments, we used unit VARISKAF-100MBS, which supports variation of main process parameters (Yurga Institute of Technology, TPU affiliate). VARISKAF-100MBS is a technological laser complex to form articles of complex shapes; it consists of ytterbium fiber laser L\u041a \u2013 100 \u2013 \u0412, scanner, PC, numerical control system, and a specially developed software tool, batch mixing powder supply system, vacuum chamber, and a gas circulation system, Fig. 1. Ytterbium fiber laser with a wavelength 1.07 \u00b5m provides output in a range 10 to 100 W. Focusing precision of a fiber laser and constant output are key factors for quality and accuracy of products \u00b10.1 mm. Laser beam manipulation with the help of a special software tool in a zone 100\u0445100\u0445100 mm allows scanning along any set contour. A diameter of a focusing laser spot is 70 \u00b5m. The surface was analyzed using a digital microscope. Drop-shaped cobalt-chromium-molybdenum powdered material \u0421\u043e28Cr3Mo with nominal dimensions of particles 90 \u00b5m is intended for SLM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure3-1.png", + "caption": "Figure 3. Plain bearing geometrical", + "texts": [], + "surrounding_texts": [ + "v b b x b y b z\nv b b x b y b z\ni i i i\nl l l l\n= + + +\n= + + +\n\n\n \n\n\n 0 1 2 3\n0 1 2 3 \n\ufeff (20)\nw c c x c y c z\nw c c x c y c z\ni i i i\nl l l l\n= + + +\n= + + +\n\n\n \n\n\n 0 1 2 3\n0 1 2 3 \n\ufeff\nBy\ufeffreplacing\ufeffthe\ufeffcoefficients\ufeffvalues,\ufeffthe\ufeffdisplacement\ufeffvector\ufeffcan\ufeffbe\ufeffwritten\ufeffas:\nP P\nu x y z u u u u v x y z v v i i j j k k l l\ni i j j k0 1\n, , , , ( ) = + + + ( ) = + + \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 v v\nw x y z w w w w k l l\ni i j j k k l l\n+\n( ) = + + +\n\n\n \n\n\n \u03b1 \u03b1 \u03b1 \u03b1 \u03b1, , \ufeff (21)\nwith\ufeff\u03b1i,\ufeff\u03b1j,\ufeff\u2026,\ufeff\u03b1l\ufeffare\ufeffthe\ufeffweighting\ufefffunction\ufefffor\ufeffthe\ufeffelement\ufeff(e). Matricially,\ufeffthe\ufeffdisplacement\ufeffof\ufeffa\ufeffpoint\ufeffamounts\ufeffto:\nu v w\ni\ni\ni\nj\nj\nj\nk\nk\n\n\n \n\n\n = \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 \u03b1 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0\n0 0 0 0\n0 0\u03b1 \u03b1 \u03b1 \u03b1\nk\nl\nl\nl\ni\ni\ni\nl\nu v w\nw\n\n\n \n\n\n \n\n\n\n \n\n\n\n \n\ufeff (22)\nu v w\nq\n\n\n \n\n\n = { }{ }\u03c6 \ufeff\nwith\ufeff \u03c6{ } \ufeffinterpolation\ufeffmatrix.\n{q} displacement Vector The\ufeffnodal\ufeffforces\ufeffequivalent\ufeffto\ufeffthe\ufeffsurface\ufeffforce\ufeff(pressure)\ufeffis\ufeffgiven\ufeffby:\nF F ds S\nt\nS\nSe\n{ } = { }\u222b\u222b \u03c6 \ufeff (23)\nThe\ufeffgeneral\ufefffundamental\ufeffequation\ufeffthat\ufefflinks\ufeffnodal\ufeffdisplacements\ufeff{q}\ufeffand\ufeffnodal\ufeffforces\ufeffis\ufeffgiven\ufeff by\ufeffthe\ufefffollowing\ufeffrelationship:\nF K q{ } = { }{ } \ufeff (24)", + "with\ufeff{K}\ufeffis\ufeffrigidity\ufeffmatrix.\nMETHodS NUMERICAL ANALySES\nModel Presentation The\ufeffnumerical\ufeffsimulation\ufeffis\ufeffcarried\ufeffout\ufefffor\ufeffa\ufeffplain\ufeffbearing\ufeffof\ufeff100\ufeffmm\ufeffdiameter\ufeffand\ufefflength\ufeffof\ufeff80\ufeffmm,\ufeff the\ufeffmain\ufeffcharacteristics\ufeffof\ufeffwhich\ufeffare\ufeffshown\ufeffin\ufeffTable\ufeff1.\ufeffThe\ufeffbearing\ufeffis\ufeffembedded\ufeffin\ufeffa\ufeffsteel\ufeffring.\ufeffThe\ufeff pad\ufeffconsists\ufeffof\ufefftwo\ufefflayers,\ufeffthe\ufefflargest\ufefflayer\ufeffis\ufeffsteel\ufeffand\ufeff38\ufeffmm\ufeffthick,\ufeffthe\ufeffinner\ufefflayer\ufeffis\ufeffa\ufefftin-based\ufeff coating\ufeff(89%)\ufeffwith\ufeffa\ufeffthickness\ufeffof\ufeff2\ufeffmm.\ufeffThis\ufeffpad\ufeffis\ufefffed\ufeffby\ufeffa\ufefffeed\ufeffgroove\ufeffwith\ufeffa\ufefflength\ufeffof\ufeff70\ufeffmm\ufeff using\ufeffthree\ufefffeed\ufeffports,\ufeffdiameter\ufeff14\ufeffmm,\ufeff(Figure\ufeff3).\nTextures Parameters Surface\ufefftexturing\ufeffis\ufeffa\ufefftechnique\ufeffused\ufeffto\ufeffimprove\ufeffthe\ufeffload\ufeffcapacity\ufeffof\ufeffvarious\ufefftribological\ufeffconjunctions,\ufeff as\ufeffwell\ufeffas\ufeffto\ufeffreduce\ufefffrictional\ufefflosses.\ufeffThe\ufefftexture\ufeffis\ufeffspherical\ufeffshape,\ufeff(Figure\ufeff4)\ufeffwith\ufeffa\ufeffdiameter\ufeffrx\ufeff", + "=\ufeff3\ufeffmm\ufeffand\ufeffthe\ufeffdepth\ufeffof\ufeffry\ufeff=\ufeff0.5\ufeffmm,\ufeffthe\ufeffaxial\ufeffdistance\ufeffbetween\ufeffthe\ufefftextures\ufeffd\ufeff=\ufeff10\ufeffmm\ufeffand\ufefftheir\ufeff angular\ufeffoffsets\ufeff\u03b1\ufeff=\ufeff10\u00b0.\nMeshing The\ufefffinite\ufeffelement\ufeffnumerical\ufeffsimulation\ufeffis\ufeffused\ufeffto\ufeffcalculate\ufeffthe\ufeffdisplacement\ufeffof\ufeffthe\ufeffinner\ufeffface\ufeffof\ufeffthe\ufeff plain\ufeffbearing.\ufeffThe\ufeffsolid\ufeffis\ufeffdecomposed\ufeffinto\ufeffa\ufeffnumber\ufeffof\ufeff4-node\ufeffor\ufeff8-node\ufefftetrahedral\ufefffinite\ufeffelements\ufeff so\ufeffthat\ufeffthese\ufeffelements\ufeffare\ufeffas\ufeffaccurate\ufeffas\ufeffpossible\ufeffin\ufeffthe\ufeffgeometry.\nThe\ufeffshaft\ufeffis\ufeffdiscretized\ufeffinto\ufeffhexahedral\ufeffelements\ufeffwith\ufeff8\ufeffnodes,\ufeff(Figure\ufeff5),\ufeff15\ufeffnodes\ufeffin\ufeffthe\ufeffaxial\ufeff direction,\ufeff54\ufeffpoints\ufeffin\ufeffcircumferential\ufeffand\ufeff22\ufeffpoints\ufeffin\ufeffthe\ufeffradial\ufeffdirection." + ] + }, + { + "image_filename": "designv11_80_0000155_fie.2018.8659280-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000155_fie.2018.8659280-Figure2-1.png", + "caption": "FIGURE 2 PROJECT ACTIVITY #1: SERIES CIRCUIT", + "texts": [ + " This kit contains electrical components are placed on the board to create basic and advanced circuits. SECTION 5: PROJECT ACTIVITIES The workshop schedule in Table I identifies five project activities. Project #1: Series & parallel circuits Following the introduction of basic circuit concepts through the variables, their units, and fundamental laws (Kirchhoff\u2019s Voltage and Current), the first project illustrates the configuration of the two basic electrical circuit configurations \u2013 series and parallel. The expected duration of the assembly and test activities in Project #1 was five minutes. Figure 2 shows the set-up of a lamp and a fan in series. The participants assembled the circuit shown in Figure 2 and were instructed not to touch the fan or motor during operation. Safety concerns form an integral part of electronic design and test and were enforced throughout this workshop. Upon placement of the fan blade on the motor (M1), and closure of the slide switch (S1), the fan will spin and the lamp (L1) should turn on. The light helps protect the motor from getting the full voltage when the slide switch is closed. The participants removed the fan and noticed how the lamp gets dimmer when the motor does not have to spin the fan blade" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002129_s12206-020-0105-8-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002129_s12206-020-0105-8-Figure1-1.png", + "caption": "Fig. 1. Case Western Reserve University's experimental set-up (A. 3 phase supply, B. Accelerometer (fan end), C. Induction motor, D. Accelerometer (drive end), E. Faulty bearing, F. Coupling, G. Torque encoder, H. Dynamometer for load, I. Data acquisition card, J. Computer).", + "texts": [ + " (3) can be modified to give the fractional order quadrature-like filter (HP) as ( ) /2 /2 , 0 0, 0 , 0 - + \u00ec > \u00ef= =\u00ed \u00ef <\u00ee jP\u03c0 jP\u03c0 e \u03c9 H \u03c9 \u03c9 e \u03c9 (4) The magnitude of a complex signal, formed from the original signal (x) and its FRHT ( \u02c6 ,)x i.e. x ,j \u02c6+ x gives a fractional envelope of the signal x. In this paper, we use only one level of fractionality - due to fractional parameter Q. Thus, the quadra- ture filter (with P = 1), as given in Eq. (3), is used. To substantiate the importance of the proposed method, we have used a standard dataset on bearing faults provided online by Case Western Reserve University [27]. The dataset contains vibration signals from a 2hp electric motor set-up as shown in Fig. 1 with 6205-2RS JEM SKF deep groove ball bearing at the drive end and 6203-2RS JEM SKF deep groove ball bearing at the fan end. These signals are acquired at the rate of 12000 or 48000 samples per second using two different accelerometers, one at the drive end and other at the fan end. Ideally, the accelerometer located near the faulty bearing should provide correct signature of fault. The accelerometer which is away from the faulty bearing also captures vibrations from other machine components. We have considered all possible permutations of bearing-accelerometer location in the cases that are represented" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001813_csci46756.2018.00091-Figure19-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001813_csci46756.2018.00091-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_80_0003733_s00158-020-02741-x-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003733_s00158-020-02741-x-Figure3-1.png", + "caption": "Fig. 3 Force analysis of the skiing process", + "texts": [ + "0014 (Gasser 2018);A is the front projection area of the athlete relative to air (m2); and v is the velocity of the athlete (m/s). The inrun is divided into three sections. The inclination of each section is different, and the geometric profiles are not the same. In order to simplify the calculation process, the athlete\u2019s force analysis and dynamic equation were established for the three sections. The inrun has a structural inclination of \u03b3. The athlete and the snowboard were the research objects, and the force analysis is shown in Fig. 3. G is the gravity of the athlete and the skateboard, FN is the support force, Fa is the air resistance, and Ff is the friction between the snowboard and the inrun. The dynamic equation of the skiing direction is m\u00feM\u00f0 \u00dea1 \u00bc m\u00feM\u00f0 \u00deg sin \u03b3\u2212\u03bc m\u00feM\u00f0 \u00de g cos \u03b3\u2212 1 2 CkA\u03c1av 2: \u00f05\u00de Then, a1 \u00bc g sin \u03b3\u2212\u03bcg cos \u03b3\u2212 CkA\u03c1av 2 2 m\u00feM\u00f0 \u00de, a1 \u00bc dv dt \u00bc ds dt dv ds \u00bc v dv ds, where s1 is the athlete skiing distance in the start zone, replacing a1. \u222bds1 \u00bc \u222b vdv g sin \u03b3\u2212\u03bcg cos \u03b3\u2212 CkA\u03c1av 2 2 m\u00feM\u00f0 \u00de \u00f06\u00de The integral was solved to obtain the relationship between the skiing velocity v1 at the end of the start zone and the related structural and environmental parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002826_joe.2019.1167-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002826_joe.2019.1167-Figure6-1.png", + "caption": "Fig. 6 The UAV modeling in design (a) The model of the UAV, (b) The structure of the UAV", + "texts": [ + " In other words, it offers the ability to accurately and efficiently simulate populations of robots in complex indoor and outdoor environments [7]. The ROS is a flexible framework for writing robot software. It is a collection of tools, libraries, and conventions that aim to simplify the task of creating complex and robust robot behavior across a wide variety of robotic platforms [8]. In ROS, we can code control functions for the UAV and run the algorithm. In the simulation environment, the model and structure of the UAV is shown in Fig. 6. Fig. 6a is the model of the UAV with the same size as the actual UAV. Fig. 6b is the assembly location of RealSense device on the UAV. The relationship between the speed of the UAV and the output of the Actor is shown in (2). The maximum speed change is 1 m/s vx = action \u22c5 x m/s vy = action \u22c5 y m/s vz = action \u22c5 z m/s (2) Fig. 7 is a simulation environment for testing. The map contains several groups of obstacles in a three-dimensional environment. For each obstacle, some necessary data are collected from all directions and locations before training. The training results are shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001569_ecce.2019.8912739-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001569_ecce.2019.8912739-Figure12-1.png", + "caption": "Fig. 12 MMS distribution, max of 207,4 MPa in the air gap bridge (red area), 148,1 MPa in the central bridge (orange area)", + "texts": [ + " The contact conditions between magnets and rotor laminations are defined at the interfaces shown in Fig.10. At these places, the rotor body and the magnets always stay in contact. No contact surfaces are configured in the other interfaces. This implies that the rotor lamination and the magnet can slide against each other. The FE distribution of the MMS in the ASRM rotor is shown in Fig.11. The peak stresses are distributed in two critical areas: locally in the air gap bridge, and globally in the central bridge; they are highlighted in Fig.12 (magnets not presented). This justifies the need for accurate analytical methods able to calculate MMS in both air gap and central bridges. Thus, we can say that the BTM outperforms the ERM in this regard. We can also notice that the stress concentration phenomenon is visible in the air gap iron bridges (Fig.12). The FE radial displacement of the ASRM rotor is also studied. The distribution of the radial displacement is reported in Fig.13 (magnets not presented). Maximum allowable displacement should not exceed 20% of the air gap length. The lowest radial displacements are located in the air gap bridges toward adjacent poles, and in the central bridges toward the rotor body. This implies zero deformation in these zones as shown in Fig.14 where both the air gap and the central bridges remain fixed. The rotor deformation caused at high speed by the centrifugal force applied on the rotor justifies the assumptions of making each pole of the rotor embedded with the adjacent poles at the air gap iron bridges and with the rotor body at the central iron bridge in the BTM (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002377_sii46433.2020.9025809-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002377_sii46433.2020.9025809-Figure3-1.png", + "caption": "Fig. 3. Skeleton of the GOOGOL GRB3016 robot with coordinate frames in the home position.", + "texts": [ + " In the update step, the state covariance matrix is updated by optimal Kalman gain K as follows: Kk+1 =Pk+1|kJ T k+1S \u22121 k+1 x\u0302k+1|k+1 = x\u0302k+1|k +Kk+1y\u0303k+1 Pk+1|k+1 =(I \u2212Kk+1Jk+1)Pk+1|k (26) where I is the identity matrix. Once the updating procedure is completed, the norm values of the state vector are calculated for every iteration. The EKF is reduced to the Newton\u2013Raphson method when Q and R are set to zero. A GOOGOL GRB3016 robot was used in this experiment to verify the proposed method. The robot can self-calibrate online in its working status. The nominal robot link parameters are shown in Table I. Fig. 3 shows the skeleton of the GOOGOL GRB3016 robot with all its coordinate frames and geometric features. The matrices R and Q for the KF can be determined by the adaptive method described in [28] and [29]. The matrices R and Q for the EKF can be calculated through the method in [30]. The GOOGOL GRB3016 robot with six DOFs requires 24 geometric parameters to be modeled. Each 3-D robot pose provides six model equations, as indicated by (19). Therefore, a unique computation of the 24 parameters requires a minimum of four pose measurements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000635_978-3-030-20377-1_1-Figure1.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000635_978-3-030-20377-1_1-Figure1.3-1.png", + "caption": "Fig. 1.3 Forces acting on a half-plane", + "texts": [ + " The vast majority of solved contact problems\u2014in the sense that there is a simple, clear algebraic expression of the contact law\u2014the relationship between the applied load and contact size\u2014and a simple description of the contact pressure distribution fall into the category of those where the bodies may be adequately idealised as half-planes. The starting point of their solution is the Flamant solution using an Airy stress function description of the stress state within a wedge (Barber 2010), particularised to the case where the internal wedge angle is 180\u25e6, Fig. 1.3. If the normal load is P and shear force Q (in each case per unit depth into the page), the appropriate Airy function is \u03c6 = \u2212r\u03b8 \u03c0 (P sin \u03b8 + Q cos \u03b8) (1.1) from which we may find the stress components by differentiation alone, using the relations \u03c3rr (r, \u03b8) = 1 r \u2202\u03c6 \u2202r + 1 r2 \u22022\u03c6 \u2202\u03b82 (1.2) \u03c3\u03b8\u03b8 (r, \u03b8) = \u22022\u03c6 \u2202r2 (1.3) \u03c3r\u03b8 (r, \u03b8) = \u2212 \u2202 \u2202r ( 1 r \u2202\u03c6 \u2202\u03b8 ) , (1.4) giving \u03c3r\u03b8 = \u03c3\u03b8\u03b8 = 0 everywhere, and So, the state of stress induced is very simple: there is only onenon-zero component, a radial direct stress acting towards the point of application of the load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003766_j.triboint.2020.106684-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003766_j.triboint.2020.106684-Figure4-1.png", + "caption": "Fig. 4. (a) The studied bearing model and (b) flow grid model of the studied bearing.", + "texts": [ + " \u03b8w is the angle between the wall and interface tangent line, which can determine the local curvature of the interface. The wall normal is given by n\u2192a = n\u2192w cos \u03b8w + t\u2192w sin \u03b8w (20) where n\u2192a is the wall normal, and n\u2192w and t\u2192w are the unit normal vector and unit tangent vector of the wall, respectively. A cylindrical roller bearing NU2326ECML which is produced by a well known bearing manufacturer in the WPG is studied. The outer diameter of bearing is 280 mm; the inner diameter of bearing is 130 mm; and the bearing has 14 rollers. The bearing model is shown in Fig. 4(a). An AOTPF model is established, and the VOF method is used to simulate the AOTPF in the bearing cavity. In Fig. 4(b), the flow grid model for bearing is shown, which is the AOTPF model. In Fig. 4(b), the model is tetrahedral unstructured meshing. The flow field mesh model has 198,032 elements. The sliding mesh model is applied at the flow edge of bearing to complete the calculation data transfer. In the proposed model, the standard atmospheric pressure is defined. The bearing end face is the pressure outlet, which is zero pressure. The Wall is a fixed boundary. In the VOF model, the air phase is compressible, which is the main phase. The viscosity of the air is 1.79 \u00d7 10\u2212 5 Pa s. The density of the air is 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002392_978-981-15-1293-3_8-Figure2.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002392_978-981-15-1293-3_8-Figure2.4-1.png", + "caption": "Fig. 2.4 Design of the pathogen-sensor platform assembled from peptide nanotubes. The peptide nanotube incorporates virus-recognition elements on the surface", + "texts": [ + " This biosensor enables a sensitive detection of glucose base on monitoring the hydrogen peroxide which is produced by an enzymatic reaction between a glucose and the glucose oxidases attached onto the peptide nanotubes. In addition, based on using ethanol dehydrogenase and NAD+, a sensitive detection of ethanol may be conducted by the marked electrocatalytic activity toward NADH.\u00a0 In another approach, non-conductive peptide nanotubes were modified with antibodies to develop sensitive sensors for viruses (Fig\u00a0 2.4) [35]. The sensor chip used the peptide nanotubes that were prepared by selfassembly from bollaamphiphilic peptide monomers and then coated with antibodies in a simple incubation process. These peptide nanotubes were assembled onto the device platform and headed to the gap between a pair of electrodes due to positive dielectrophoresis. Pathogen detection was happened through the difference in the dielectric properties of viral particles and water molecules. At the molecular level, biological processes are based on physical and chemical conditions whose fundamental chemical systematics in the Periodic System of the Elements (PSE) were introduced by Mendeleyer and Meyer in 1869" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002546_cyber46603.2019.9066743-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002546_cyber46603.2019.9066743-Figure6-1.png", + "caption": "Figure 6. Trajectories of the robots with a faulty robot", + "texts": [ + " From it, the flocks have reached the position of maximum value in the field can be observed, and the velocity of each robot has decreased to zero because seeking has been completed. This result indicates that the proposed DSSF scheme is effective for source-seeking. Fig. 5 demonstrates the velocities of all the robots varying with time in the process of source-seeking. It can observe that robot velocities gradually reach the similar value due to flocking, and after the source of the field being found, the velocities of all the robots decrease to zero. Fig. 6 demonstrates the trajectories of the robots with assuming that a fault robot exists. From this figure, it can find that all the robots initially maintain flocking. Then, the flocking is breaked by a fault robot, and is formed again. Finally, other robots successfully locate the source. Fig. 7 demonstrates the velocities of the robots with considering that a fault robot exists. From Fig. 7, a fault robot occur at time 3500 ms. After a short confusion time, other robots reach a similar velocity again, i.e. they form a flock again. The phenomena in Fig. 6 and Fig. 7 indicate that the proposed DSSF scheme enables that the robot flocking has the survivability of tolerating fault robots. In this paper, the problem of source-seeking in environmental field is considered. A DSSF scheme to resolve this problem is proposed. With the DSSF scheme, the seeking velocities of robots can be accurately decided through a two-level judgment processes, and the source of a field can be successfully located even when there is a fault robot. The proposed DSSF scheme is simulated and the results demonstrate the effectiveness and robustness of the DSSF scheme" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000374_978-94-024-1620-6_11-Figure11.12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000374_978-94-024-1620-6_11-Figure11.12-1.png", + "caption": "Fig. 11.12 Illustration of the molecular structure of non-stacked, grafted reduce graphene oxide with HFHPB derivative uptake of DMMP. (Reprinted with permission from [62]. Copyright 2016 Nature Publishing Group)", + "texts": [ + " These transducers work by measuring change of frequency in a quartz crystal resonator when mass variations per unit area occur, (e.g., mass changing before and after analyte capture). A novel gravimetric sensor example can be found in work reported by Hwang and coworkers [62]. Based on reduced non-stacked graphene oxide (NSrGO)hexfluorohydoroxypropanyl benzene (HFHPB), the sensor allowed the detection of dimethyl methyl phosphonate (DMMP), used as simulant of sarin gas. They reported the synthesis of 3D nanostructured porous NSrGO-HFHPB in which HFHPB worked as receptor for effectively detection of DMMP (Fig. 11.12). The HFHPB receptor is capable to form hydrogen bonding with the phosphonate group present in most nerve agents. The average pore size of composite was expected to be large enough for an easy diffusion of DMMP molecules through matrix, providing a fast response and high sensitivity due to the large exposed surface area. For evaluating performance of sensor, authors explored sensing behavior of GO, NSrGO and the composite hybridized with HFHPB receptors finding that NSrGOHFHPB film sensor had 13 times greater sensitivity then the other composites" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001707_icems.2019.8921756-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001707_icems.2019.8921756-Figure2-1.png", + "caption": "Fig. 2 (a)-(c) shows the second mode shape to the fourth mode shape of the stator core, which corresponds to the shape of radial vibration.", + "texts": [], + "surrounding_texts": [ + "The modal analysis of the stator had closed relevance to the structure parameters. Different sizes of the stator core can affect the trend of the natural frequencies, therefore, we can study the stator structure to seek for what factors which may have influences on the natural frequencies. There are two main factors affecting the natural frequency of the stator core: mass and stiffness. The influence of tooth parameters on mass is easy to analyze, but the influence of tooth parameters on stiffness is hard to get. In order to analyze the influence of tooth parameters on stiffness, mass is kept unchanged by changing the density of the stator. Change the parameters of the stator tooth and get the relationship between parameters and natural frequencies. The tooth parameters are: tooth width and tooth depth and number of slots. The FEM results are shown in Fig.3. In case the mass remains the same, it is clear that the stiffness is proportional to the number of slots and tooth width, and inversely proportional to the tooth depth. III. NATURAL FREQUENCY ANALYTICAL METHOD There are generally three methods for analyzing the natural frequency of a motor: FEM\u3001AM and hammer test. FEM could model the motor accurately, AM could get natural frequency faster and simpler and get analytical formula. Hammer test could verify the correctness of the results from an experimental perspective. This part presents a more accurate AM for calculating the natural frequencies of a 36-slot 6-pole IPMSM. The natural frequencies of the stator with different tooth depth, tooth width, yoke thickness and axial length are calculated by AM and compared with FEM. The following parts are the results of the analytical method." + ] + }, + { + "image_filename": "designv11_80_0003452_0954405420951101-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003452_0954405420951101-Figure2-1.png", + "caption": "Figure 2. Comparison of standard transition curve and elliptical curve.", + "texts": [ + " The geometry of the proposed gear tooth profile The structure of the gear rolling system with axialinfeed developed by SKLMT at Chongqing University is shown in Figure 1. Two rolling tools have the same rotational speed controlled by motors and transmission system, which consists of belt pulleys and worm gears. The racial-infeed cylinder adjusts the center distance according to different workpiece. The workpiece are pushed through the space between two rolling gears by axial-infeed cylinder and the workpiece is deformed into a gear with involute teeth through the meshing with rolling gears. The standard gear profile and optimized elliptical curve are shown in Figure 2. The definition of all symbols is shown in the Table 1. A coordinate system is established on the gear center. The center of designed elliptical curve is noted as C1 and C1 is located on the symmetry axis. a and b are radii of the elliptical curve on the semi-major axis and semi-minor axis respectively. d is noted as the distance between the center of elliptical curve and pitch circle. Assumed point M is on the elliptical curve, the parametric equation of the circular fillet can be expressed by function of u1: x1m = a3 cos u1 y1m = b3 sin u1 (r1 d) \u00f01\u00de The pitch radius is denoted as r1 and the pressure angle on the pitch circle is 20 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002433_j.promfg.2019.07.038-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002433_j.promfg.2019.07.038-Figure1-1.png", + "caption": "Fig. 1. (a) Test specimens; (b) Fixturing device to hold the specimens; (c) Conoscan 4000 3D scanner", + "texts": [ + " In order to analyse the influence of the material colour and the capability of the CH sensor for measuring different layer thicknesses, four rectangular PLA specimens of different colour (white, blue, red and black) were manufactured using a Sigma BCN3D machine. These four colours represent different tonalities compared to which the wavelength of the laser CH sensor may have different behaviour. Each of them contains four circular surfaces of 15 mm diameter, in positions distant 18 mm from each other, with different height achieved by deposition of 1, 2, 3 and 4 layers of 0.1 mm thickness (Fig. 1a), respectively. The PLA specimens were attached to aluminium plates to increase their stiffness and to facilitate their orientation and positioning in the fixture device of the digitizing machine. The digitizing tests carried out in this study were performed with an Optimet\u2122 ConoPoint-10 point-type conoscopic holography sensor integrated in the Optimet Conoscan 4000 3D high precision scanner (Fig. 1c). The G. Vali\u00f1o et al. / Procedia Manufacturing 41 (2019) 129\u2013136 131 Author name / Procedia Manufacturing 00 (2020) 000\u2013000 3 visible light source of the sensor is a Class II laser diode which wavelength is 655 nm. A lens of 25 mm focal length and 1.8 mm depth of field (DOF) was connected to the sensor. Each sensor reading provides the value of distance between the transmitter and the projection of the laser beam on the material surface (spot). According to the Z scanner stroke and the sensor\u2019s DOF, a specific fixture was made to hold the specimens within the machine working area and ensure tests repeatability (Fig. 1b). Two main parameters are necessary to adjust the CH sensor (Table 1): Power Level (P), which represents the value of the laser beam energy required. Working Frequency (F), which represents the data acquisition rate. For a given F, the value of P should be adjusted so that enough light is received by the sensor. For a low P level, the amount of light reflected off the surface could not be sufficient and quality of the measurement would drop in consequence. On the other hand, a high P level may yield a saturated signal and measurements would not be reliable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003110_s40194-020-00955-7-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003110_s40194-020-00955-7-Figure2-1.png", + "caption": "Fig. 2 Physical model for each parameter adjustment. (a) Defocusing distance adjustment. (b) Powder feeding angle adjustment. (c) Pool width adjustment", + "texts": [ + " (20) in [13] and Eq. (3) in [8]). In this model, the outflow and injection positions of central particles are changed with adjustments of defocusing distance and powder feeding angle, and the molten pool boundary moves with variation of the pool width. The new coordinate position information can be re-substituted into the original theoretical model to calculate the powder transport ratio after adjustment of injection parameters. \u2013 The effect of defocusing distance adjustment on the physical model is presented in Fig. 2(a). The adjustment amount is \u03c7 and the movement of the powder nozzle away from the workpiece is taken as positive. By +\u03c7 adjustment, therefore, the height of the outflow position and the range of powder flow boundaries (P1, P2) increase, and the injection position of the central particles moves in +X direction. The situation is reverse for \u2212\u03c7 adjustment. A parallel laser beam is assumed, so the width of the molten pool would not be changed with the adjustment. \u2013 The effect of the powder feeding angle adjustment on the physical model is presented in Fig. 2(b). As the nozzle orifice position is fixed, the outflow position is changed by the rotation around the fixed point of tube (fulcrum). The adjustment amount is \u03b1 and rotation of powder nozzle in clockwise direction is taken as positive. By +\u03b1 adjustment, injection position and powder flow boundaries (P1, P2) move toward \u2212X direction. The situation is reverse by \u2212\u03b1 adjustment. Here we assume that the molten pool width would not be changed with the variation of particle concentration distribution caused by this adjustment. \u2013 The effect of molten pool width adjustment on the physical model is presented in Fig. 2(c). The adjustment amount is \u03b4. Boundaries (P3, P4) of powders falling into the pool are moved by this adjustment, and the powder transport ratio is changed. In order to make these three adjustments unified into one mathematical model, the sequence of the derivation process is defocusing distance +\u03c7, powder feeding angle +\u03b1, and pool width \u2212\u03b4. Based on the model analysis above, the key of modeling is information on outflow position, injection position, and pool boundary. The trajectory of the central particle from the outflow to injection position is related with +\u03c7 and +\u03b1 adjustments, which are hard to determine", + " \u00f019\u00de The outflow position vector OA2 ! is expressed as OA2 ! \u00bc OA1 !\u00fe A1A2 ! \u00bc OA1 !\u00fe A1M !\u00feMA2 ! \u00bc \u2212W\u2212 L\u00fe \u03c7\u00f0 \u00de=tan\u03b3\u2212 MA2 ! cos\u03b3 MA2 ! i\u00fe MA2 ! sin\u03b3 MA2 ! j \u00f020\u00de The injection offset vector ON ! can be given by Eq. (21), where x1, x2 are calculated in Eqs. (6, 18). ON ! \u00bc OM !\u00feMN ! \u00bc x1\u2212x2\u00f0 \u00de i \u00f021\u00de The boundary condition of powders falling into the pool is only changed with the adjustment (\u03b4) of the molten pool width, by which the particle concentration distribution is not affected. As depicted in Fig. 2(c), the boundary after \u2212\u03b4 adjustment is determined by Eq. (22). x P3\u00f0 \u00de \u00bc D\u2212\u03b4\u00f0 \u00de=2 x P4\u00f0 \u00de \u00bc \u2212 D\u2212\u03b4\u00f0 \u00de=2 \u00f022\u00de As confirmed, the powder flux actually tends to be of a Gaussian distribution [8], and the key position information are brought into the original Gaussian model merely to calculate powder transport ratio. Information of each key position and vector parameter after adjustments is listed in Table 1. Figure 1 reflects the state of the powder flow field before adjustment of the powder injection parameter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001455_012165-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001455_012165-Figure1-1.png", + "caption": "Figure 1. Von Mises Stress ( Max: 43.55 MPa)", + "texts": [], + "surrounding_texts": [ + "IOP Conf. Series: Earth and Environmental Science 343 (2019) 012165 IOP Publishing doi:10.1088/1755-1315/343/1/012165" + ] + }, + { + "image_filename": "designv11_80_0002057_s00006-019-1039-z-Figure20-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002057_s00006-019-1039-z-Figure20-1.png", + "caption": "Figure 20. a A 6R serial manipulator. b The network representation of a 6R serial manipulator in terms of enumerated kinematic pairs", + "texts": [ + " (43f)\u2013(43h), for i = (1, 2, 3) B1 2 = V {2, 3} \u03b46\u03b411 2 (43c) B2 2 = V {1, 3} \u03b41\u03b411 2 (43d) B3 2 = V {1, 2} \u03b41\u03b46 2 (43e) C1 2 = L1 \u2227( L2 \u03b46 2 + L3 \u03b411 2 ) (43f) C2 2 = L2 \u2227( L1 \u03b41 2 + L3 \u03b411 2 ) (43g) C3 2 = L3 \u2227( L1 \u03b41 2 + L2 \u03b46 2 ) (43h) As can be seen in the reverse statics and kinematics case study plots (Figs. 18a\u2013c and 19a\u2013c) and in Table 2 below, there is a strong correlation between the reverse hypervolume functions and their statics and kinematics counterparts. The hypervolume magnitude plots are geometrically similar to the input solutions of parallel statics and kinematics problems. 3.3.1. A 6R Serial Manipulator Forward Statics and Kinematics Comparison. In this example, the 6R manipulator\u2019s (Fig. 20) statics and kinematics daughter hypervolumes for the [e1, e2, e3 ] bases, FV (s,f) [e1, e2, e3 ] and FV (s,k) [e1, e2, e3 ], (from Eqs. (30a) and (23a) respectively) are compared against the forward static and kinematic end effector outputs of |fE | and |vE | respectively. Where FV (s,k) = X \u03b8 1,7\u03b31 \u2227 X \u03b8 2,7\u03b32 \u2227 X \u03b8 3,7\u03b33 \u2227 X \u03b8 47\u03b34 \u2227 X \u03b8 5,7\u03b35 \u2227 X \u03b8 6,7\u03b36 (44a) and FV (s,f) = Min[V \u22121 1 7\u03b41,V \u22121 2 7\u03b42,V \u22121 3 7\u03b43,V \u22121 4 7\u03b44,V \u22121 5 7\u03b45,V \u22121 6 7\u03b46] (44b) where V \u22121 i n+1 is expanded according to Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001763_aeat-04-2019-0087-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001763_aeat-04-2019-0087-Figure11-1.png", + "caption": "Figure 11 A quadrotor testbed hosting four edge controllers, mounted at the tips of the two 600mm airframe beams", + "texts": [ + " Another possible explanation is that the edge controller uses superior commutation sensing circuitry, lower on-state resistance transistors and betterMOSFET gate drivers among others. The edge controller working in 2S mode has efficiency advantages that result from the lower input voltage, but this benefit is related to the physics of the setup, not to the hardware or software implemented in the edge controller (Zabunov andMardirossian, 2019). The next test experiment was conveyed on a testbed platform shown in Figure 11. This is an X-framed quadcopter encompassing four 100W motors rotating 10 in. propellers. Each motor is controlled by a dedicated edge controller. The controllers are mounted at the tips of the airframe beams in close proximity to the motors. A Z-pilot nano autopilot is used, visible in the center of the airframe (Zabunov, 2019). The communication between the autopilot and the edge controllers is through I2C interface. Due to the very short distances nonshielded cables were used. The multirotor is powered by 6 Li-ion 18650 cells with 3400 mAh capacity each" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure8.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure8.6-1.png", + "caption": "Fig. 8.6 Notch root mean stress sensitivity to load history. Load sequences (a) and (b) are equivalent, but the overload cycle\u2019s peak and valley are interchanged. This causes a dramatic change in local mean stress at the fatigue notch root as seen in (c) (GH vs. DE). This effect is modelled in engineering notch fatigue analysis to account for the effect of load history on fatigue life. The effect ought to vanish under fully elastic response. But in reality it actually gets more accentuated! Research at BISS was able to explain why this may be so", + "texts": [ + " Nevertheless, the LDA concept has been widely applied to engineering structures and is still remarkably popular to the present day. Most fatigue failures occur at notches. When the applied stress multiplied by the elastic stress concentration factor, Kt, exceeds the yield stress, the notch root will see inelastic response. Given that stress-strain hysteresis occurs, the notch root mean stress will not only be different from the applied mean stress, but also become load sequence dependent, as illustrated in Fig. 8.6. This raised the possibility that provided one could compute the actual notch root mean stress for individual cycles of a given load sequence, then maybe the LDA 112 R. Sunder concept would enable realistic estimates of notch fatigue life. Indeed, the Local Stress Strain (LSS) approach that performs such modelling has served industrial design for over 40 years. However, the LSS approach fails the very basic test of fully-elastic notch root response. This problem is important, because designers try to keep local stresses elastic: but in this case the local mean stress would be rendered insensitive to the load sequence", + " Since only surface atomic layers are affected, the crack-tip surface layer effect (and hence the near-tip MSE) decreases with increasing growth rate. In other words, our second and third experiments essentially demonstrated that whilst mechanisms such as closure affect the mechanics (driving force) of fatigue, near-tip mean stress affects the material\u2019s resistance to environmental FCG. This result is explained by the BMF theory. An important corollary of the BMF theory relates to the consequence of near-tip hysteretic cyclic stress-strain response. As is seen from Fig. 8.6, hysteretic response makes load interactions cycle-sequence sensitive, even if there is no crack extension. Thus hysteretic-response-induced changes of the near-tip mean stress can by themselves change the cycle-by-cycle resistance to BMF. In contrast, crack closure cannot exhibit cycle sequence sensitivity because the crack wake can yield only in compression. This was shown by a fourth experiment, discussed next. Fourth experiment: Steps of extremely small load amplitudes superimposed on the rising and falling halves of periodic overloads highlighted the hysteretic nature of variable-amplitude near-threshold fatigue crack growth, with cycle sequence sensitive FCG rates varying by over an order of magnitude (Sunder 2005)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000726_978-3-030-21013-7_15-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000726_978-3-030-21013-7_15-Figure1-1.png", + "caption": "Fig. 1 The geometric and modal parameters of the flexible tool-rigid workpiece model", + "texts": [ + " In the model, w is the depth of cut, \u03b8 j (t) = \u03c0 t/30 + 2\u03c0( j \u2212 1)/N is the angular displacement of the jth cutting edges (that is, j = 1,2,\u2026,N), g j (t) = 0.5 ( 1 + sgn ( sin ( \u03b8 j (t) \u2212 tan\u22121 P) \u2212 sin ( \u03b8s \u2212 tan\u22121P))) is the screening function where P = (sin \u03b8s \u2212 sin \u03b8e)/(cos \u03b8s \u2212 cos \u03b8e), \u03b8s and \u03b8e are the start and end angles of the cutting interval, B is the radial depth of cut and D is the tool diameter. For upmilling, \u03b8s = 0 and \u03b8e = cos\u22121(1 \u2212 2\u03c1)while for down-milling, \u03b8s = cos\u22121(2\u03c1 \u2212 1) and \u03b8e = \u03c0 where \u03c1 = B/D is the radial immersion. A illustration of this model showing the geometric and modal parameters is shown in Fig. 1. On inserting the above model matrices in the constructed monodromy matrix and substituting the numerical values given in Table 1, the stability diagrams given in Fig. 2 are computed. The results agree with the known results in [4, 23]. This milling case is illustrated in Fig. 3 showing the primary motion (spindle rotation ) and the secondary motion (feed v) of the tool. The thin-walled workpiece is much more flexible than the tool and it is thus considered compliant while the tool is considered rigid" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001674_icems.2019.8921768-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001674_icems.2019.8921768-Figure6-1.png", + "caption": "Fig. 6. Fluid-solid coupling model of the induction motor", + "texts": [ + " 5 shows the variation of current and torque against with time when the motor is suddenly loaded with periodic load in the operating process. The period and the duration of the added periodic load is 4 and 2s, respectively. The current and electromagnetic torque will oscillate dynamically in each load period when the load is periodically varied. III. TEMPERATURE RISE The ventilation system structure of the motor consisting of axial and radial ventilation ducts. A 3D fluid-solid coupling model of the motor is built as well for calculating the temperature rise of the motor under different load models. Fig. 6 shows the 3D coupling model, where S1, S2 and S3 represent three different fluid inlets located at the stator back and rotor field spider, respectively. The transient heat source and speed boundary of the proposed 3D model are determined based on the transient current, the impedance and the speed. The air volume inside the motor is obtained by the equivalent airflow consisting of the wind resistance and the wind pressure source [7]. According to the CFD theory, the temperature rise of the motor running under 1, 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002330_tpec48276.2020.9042583-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002330_tpec48276.2020.9042583-Figure2-1.png", + "caption": "Fig. 2. Meshed FE model of a healthy BLDC motor (b) Magnetic density BM profile of the PMs.", + "texts": [ + " ( ) = \u23a9\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a7 1 0 \u2264 <1 \u2212 \u2212 \u2264 <\u22121 \u2264 <\u22121 + \u2212 \u2264 < 2 (6) The closed loop BLDC motor drive is developed using the parameters as given in Table.1-2 and an equivalent block diagram is given in Fig. 1. while the ideal trapezoidal backEMF obtained from (6) is given in Fig. 1. The hybrid analytical-numerical model of a BLDC motor is developed using the parameters as given from Table-I-II, through the replacement of an ideal back-EMF profile (5) in the analytical model with the numerically obtained BM plots (excluding slotting effect) developed using FEMM as shown in Fig. 2. For the calculation of EB, the motor back-EMF constant (KE) obtained from the machine rating, is multiplied with the shaping function and mechanical speed of the motor. As given in (5), KE is multiplied with the obtained BM and the motor\u2019s speed to give the required back-EMF. More elaborated modeling details are given by the authors in [17-18]. Authorized licensed use limited to: University of Exeter. Downloaded on May 07,2020 at 15:00:17 UTC from IEEE Xplore. Restrictions apply. For the KE =0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003517_1464419320955114-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003517_1464419320955114-Figure2-1.png", + "caption": "Figure 2. Bearing construction (a) 3D cut section of bearing, (b) elliptical contact area representation, (c) spring mass model of bearing.", + "texts": [ + " In the present study, validation of a mathematical model has performed with the help of systematic experiment. An adequate correlation between theoretical and experimental results has presented in this research work. Schematic Diagram of Induction motor of 60 kW capacity is shown in Figure 1. The deep groove ball bearing is mounted on a drive and non-drive end position. The major focus of this study is Non-drive end side bearing, which creates huge disturbance if pre-dispatch handling damage occurs. Free body diagram with coordinate system X and Y direction are shown in Figure 2(b) and (c). During design of the mathematical model, the following realistic assump- tions are considered:- \u2022 The housing, bearing and shaft are modelled using three DOF system. \u2022 Forces due to Centrifugal phenomenon on the balls are negligible. \u2022 The mass of races is incorporated with shaft and housing. \u2022 Skidding movements of rolling elements are neglected. \u2022 Rolling elements are uniformly positioned across the circumference of inner raceway with the help of retainer. \u2022 Contact formation between ball and races consid- ered as non-Linear hertzian contact deformation", + " In this article, the actual waviness defect is considered from pre-operational damage of bearing races as shown in Figure 3. Generally, Pre-operational damage is occurred during ball and races assembly as well as during induction motor-bearing assembly. Outer ring\u2019s round shape is distorted because the ring gets compressive loads during these pre-operations. Whenever the position of the ball and defect peak encounters, these waviness points are incorporated in the mathematical model (equation (8)) with maximum peak waviness of 2.5 mm. In Figure 2(b), the elliptical contact area is presented and deformation of the contact between the ball and track raceway can be calculated by using the hertzian load-deformation and it is expressed as follows:28 Q \u00bc Kd 3 2 (1) Deflection in radial direction at rotating element is written as follows: d/ \u00bc drcos/ 1 2 Pd (2) At time t, the angular position /i at is expressed as: /i \u00bc 2pi Nb \u00fe xct\u00fe /0 (3) Where the cage speed is mentioned as: xc \u00bc 1 d D xs 2 (4) As per internal design geometry of the bearing, the applied load on the contact of ball and track raceway can be calculated at position / with the help of following expression: Q/ \u00bc Qmax 1 1 2e 1 cos/\u00f0 \u00de (5) Where e \u00bc 1 2 1 Pd 2dr The summation of horizontal components accumulated at contact can be expressed mathematically as follows W \u00bc X/\u00bc w1 /\u00bc0 Q/cos/ (6) The deflection in the radial direction is evaluated by the following expression di \u00bc Xs Xb\u00f0 \u00decos/i \u00fe Ys Yb\u00f0 \u00desin/i c (7) Where diametrical clearance is (c) \u00bcPdl2\u00f01 cos1\u00de By considering that Xh; Yh and Zh as the deflections of the housing in three direction, the deflection of the ball in the radial direction is expressed as di \u00bc Xs Xb\u00f0 \u00decos/i \u00fe Ys Yb\u00f0 \u00desin/i \u00fe Xb Xh\u00f0 \u00decos/i \u00fe Yb Yh\u00f0 \u00desin/i \u00fe Zs Zb\u00f0 \u00decos/i \u00fe Zb Zh\u00f0 \u00decos/i c (8) Stiffness K for both ring and Damping coefficients of balls due to oil film is expressed here Kinner i \u00bc 2 ffiffiffi 2 p E 1 v2 3 P qinner\u00de 1=2 \u00f0d inner\u00de 1:5 KOuter o \u00bc 2 ffiffiffi 2 p E 1 v2 3 P qOuter\u00de 1=2 \u00f0d Outer\u00de 1:5 cb \u00bc 3ga4h 2h30 The total damping coefficient c is, { \u00bc Xu1 u\u00bc u1 {u cosu (9) Where, {u \u00bc {1{0 {i \u00fe {o The combined mass of outer ring and housing can be mentioned as m2 \u00bc mOR \u00femh (10) For multi degree freedom, the equations of motion can be presented as M\u00bd f\u20acxg \u00fe \u00bdC f _xg \u00fe K\u00bd xf g \u00bc fF t\u00f0 \u00deg (11) By considering, the all elements in mathematical model equation (11) is presented as shown in equation (12): m1 k2 0 k2 m2 k3 0 k3 m3 2 4 3 5 \u20acx1 \u20acx2 \u20acx3 8< : 9= ; \u00fe c c 0 c c 0 0 0 0 2 4 3 5 _x1 _x2 _x3 8< : 9= ; \u00fe k1 \u00fe k2 k2 0 k2 k2 \u00fe k3 k3 0 k3 k3 2 4 3 5 x1 x2 x3 8< : 9= ; \u00bc F1 t\u00f0 \u00de F2 t\u00f0 \u00de F3 t\u00f0 \u00de 8< : 9= ; (12) However, {M} is the Mass matrix, {C} is the damping coefficient and {K} is the stiffness matrix of bearing system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000457_sdtp.12989-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000457_sdtp.12989-Figure9-1.png", + "caption": "Figure 9. Structure of flexible VA-mode NPS-LCDs in (a) off state (Dark), and (b) on state (Bright).", + "texts": [ + " This light scattering results in the degradation of contrast ratio and transmittance of LCDs. In this research, we propose to use the nano-phase-separated LCs to control the viscoelasticity and alignment of LCs to improve a flexibility, contrast ratio and transmittance of sheet-type LCDs. Nano-phase-separated (NPS) LCs has a nano-size polymer network in LCs which is formed by the polymerization induced phase separation process. Due to the vertically aligned nano-size network structure, NPS-LCs can achieve uniform LC alignment and high contrast ratio without light scattering (Fig. 9). In addition, the fast decay switching of 1 ms was achieved by enhancing an anchoring force at the boundary between LCs and polymer networks [9]. In the NPS-LCs, the maximum transmittance tends to drop by the slight light scattering due to the variations of azimuthal direction in tilted liquid crystals [10]. Therefore, we investigated a fabrication process of NPS-LCs to achieve a high transmittance without light scattering comparable to the conventional VA-mode LCDs. We used collimated UV light source instead of the conventional diffused UV light to control the uniformity of alignment of nano-sized polymer network" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000746_s40722-019-00139-6-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000746_s40722-019-00139-6-Figure3-1.png", + "caption": "Fig. 3 Structure of T-foil with flap", + "texts": [], + "surrounding_texts": [ + "T-foils are significant in improving the seakeeping performance of WPCs. The T-foil is small relative to the ship\u2019s hull, but\u2014with a reasonable installation\u2014increases the damping of the ship and its resistance to wave-induced forces and moments and reduces vertical motions. The vertical acceleration is usually produced by heave and pitch and affects the seaworthiness of the ship. Tables\u00a02 and 3 show the principle parameters of the T-foil and trim tab, respectively; Figs.\u00a03 and 4 show the structure of the actuators (von Sicard 2002). The motion of the trim tabs is limited upward, while the fin of the T-foil can move freely upward and downward. The design of the actuators was based on the dimensions of the ship and hydraulic cylinder capabilities, and the relevant physics dictated the boundaries. The T-foil fins were trapezoidal with a 3-m span, 2.5\u00a0maximum chord, 13.5\u00b0/s maximum rotational speed, and \u00b1 15\u00b0 maximum angle. The fin flaps also had a 13.5\u00b0/s maximum rotational speed and \u00b1 15\u00b0 maximum angle. The trim tabs were rectangular with a 4.8-m span, 1.1 chord, 13.5\u00b0/s maximum rotational speed, and \u2212 15\u00b0 maximum angle. The actuators offer lift force at the price of drag and other degrading phenomena, such as cavitation and turbulence. As mentioned, the T-foil can produce both upward and downward forces, while the trim tab can only generate an upward force. As such, the T-foil plays the leading role in anti-vertical control, while the trim tab plays a supplementary role in anti-vertical control. The lift force is a function of the actuator angle \u03b1 and the fluid speed U given by (Esteban et\u00a0al. 2005) where lT and lF are the locations of the T-foil and trim tab on the ship, respectively; CLT and CLF are the lift coefficients of the T-foil and trim tab, respectively; \u03b1T is the T-foil fin angle of attack; \u03b1F is the trim tab angle of attack; AT is the area of the T-foil; AF is the area of trim tab; \u03c1 is the fluid density; and U is the ship velocity. From the experiment, the relationships between the fin angles and the T-foil/trim tab lift coefficients are shown in Fig.\u00a05. A sketch of the WPC with T-foils and trim tabs is shown in Fig.\u00a06. (3) FT = 1 2 (AT + ATF)CLT ( T , TF ) U2 FF = 1 2 AFCLF( F)U 2 MT = 1 2 lT ATCLT ( T ) U2 FF = 1 2 lF AFCLF( F)U 2, The T-foil lift coefficient, CLT, is a two-variable function of \u03b1T and \u03b1TF and can be obtained by a regression analysis (Thomas 1998) using experimental data. The stability of the WPC with RCS is herein defined as: after the WPC is disturbed by external or ride control forces, it deviates from its original equilibrium position; when the same force is canceled, the boat has the ability to return to its original equilibrium state. We obtained an\u00a0inequality with data in this design: where \u03b1Fmax, \u03b1Tmax, and \u03be5max are the maximum angles of the T-foil, trim tab, and pitch motion, respectively. This means that the maximum moment generated by the RCS cannot exceed the restore moment of the ship with a maximum pitch angle\u2014that is, the RCS will not affect the stability of the hull." + ] + }, + { + "image_filename": "designv11_80_0000020_mwscas.2018.8623843-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000020_mwscas.2018.8623843-Figure4-1.png", + "caption": "Fig. 4. Octorotor with standard configurations", + "texts": [], + "surrounding_texts": [ + "In this paper, we assume that all rotors are the same, distributed evenly and coplanar, and the distance from each rotor to the geometric center of the multirotor \u2113 is equal [3]. The multirotor have standard symmetrical configurations that the clockwise rotating rotor is adjacent to the counterclockwise rotating rotor as shown in Fig. 2\u20134. Fi, (i = 1, 2, . . . , 2p, p = 2, 3, 4) and Mi, (i = 1, 2, . . . , 2p, p = 2, 3, 4) of Figure 2\u20134, represent vertical force and moment, respectively. Each motor of multirotors has an angular speed \u03c9i and produces a vertical force Fi according to: Fi = kFi\u03c9 2 Mi, i = 1, 2, . . . , 2p, p = 2, 3, 4. (7) Each motor also produces a moment according to: Mi = kMi\u03c9 2 Mi, i = 1, 2, . . . , 2p, p = 2, 3, 4. (8) In practice, simple lumped parameter models are applied, such that kF > 0 and kM > 0 are given as constants that can be easily determined from static thrust tests. For multirotors from quadrotors: p = 2, hexarotors: p = 3 until octorotors: p = 4 that have standard symmetrical configurations, we define the moments of Eq. (3): Frot(u), and the translational forces of Eq. (6): Ftra(u) as Srot2pu2p, and Stra2pu2p, p = 2, 3, 4, respectively. For quadrotors: p = 2, Srot4 = 0 \u2212\u2113 \u00b7 kF2 0 \u2113 \u00b7 kF4 \u2113 \u00b7 kF1 0 \u2212\u2113 \u00b7 kF3 0 \u2212kM1 kM2 \u2212kM3 kM4 , (9) u4 = 4\u2211 i=1 \u03c92 Mi\u03b5Mi. (10) For hexarotors: p = 3, Srot6 = 0 \u2212 \u221a 3 2 \u2113 \u00b7 kF2 \u2212 \u221a 3 2 \u2113 \u00b7 kF3 0 \u221a 3 2 \u2113 \u00b7 kF5 \u221a 3 2 \u2113 \u00b7 kF6 \u2113 \u00b7 kF1 0.5\u2113 \u00b7 kF2 \u22120.5\u2113 \u00b7 kF3 \u2212\u2113 \u00b7 kF4 \u22120.5\u2113 \u00b7 kF5 0.5\u2113 \u00b7 kF6 \u2212kM1 kM2 \u2212kM3 kM4 \u2212kM5 kM6 ,(11) u6 = 6\u2211 i=1 \u03c92 Mi\u03b5Mi. (12) For octorotors: p = 4, Srot8 = 0 \u2212 \u221a 2 2 \u2113 \u00b7 kF2 \u2212\u2113 \u00b7 kF3 \u2212 \u221a 2 2 \u2113 \u00b7 kF4 0 \u221a 2 2 \u2113 \u00b7 kF6 \u2113 \u00b7 kF7 \u221a 2 2 \u2113 \u00b7 kF8 \u2113 \u00b7 kF1 \u221a 2 2 \u2113 \u00b7 kF2 0 \u2212 \u221a 2 2 \u2113 \u00b7 kF4 \u2212\u2113 \u00b7 kF5 \u2212 \u221a 2 2 \u2113 \u00b7 kF6 0 \u221a 2 2 \u2113 \u00b7 kF8 \u2212kM1 kM2 \u2212kM3 kM4 \u2212kM5 kM6 \u2212kM7 kM8 ,(13) u8 = 8\u2211 i=1 \u03c92 Mi\u03b5Mi. (14) For multirotors: p = 2, 3, 4, Stra2p = 0 0 \u00b7 \u00b7 \u00b7 0 0 0 \u00b7 \u00b7 \u00b7 0 kF1 kF2 \u00b7 \u00b7 \u00b7 kF2p , (15) u2p = 2p\u2211 i=1 \u03c92 Mi\u03b5Mi, (16) where span{\u03b5M1, \u03b5M2, . . . , \u03b5M2p}: the base vectors of a right 2p-dimensional real vector space R2p. Here we summarize the state equations of the multirotor with the symmetrical standard configurations: p = 2, 3, 4, as follows: d dt ( x x\u0307 ) = ( x\u0307 Y(\u03b7, x\u0307) + Z(\u03b7)Srot2pu2p ) , (17) d dt ( r r\u0307 ) = ( r\u0307 \u2212ge3 + 1 m B(\u03d51(t, (x0, x\u03070)T,u2p))Stra2pu2p ) , (18) where u2p \u2208 \u039b2p \u2282 R2p. III. Multirotor flight states, or operating points This section provides Definition 1 of flight states, or the operating points. Definition 1: Operating points (or all flight states) of multirotors: (xop, x\u0307op)T \u2208 R3 \u00d7 R3, xop = (\u03c8op, \u03b7op)T, \u03b7op = (\u03b8op, \u03d5op)T, and (rop, r\u0307op)T \u2208 R3 \u00d7 R3 are determined by( x\u0307op Z(\u03b7op)Srot2pu2p(op) ) = ( 03 03 ) , (19)( r\u0307op \u27e8e3,\u2212ge3 + 1 m B(xop)Stra2pu2p(op)\u27e9 ) = ( 03 c ) , (20) Z(\u03b7op) = (B(xop)I\u0302B(xop)T \u00b7 \u03c9x\u0307(\u03c8op, \u03b8op))\u22121 \u00b7 B(xop), (21) r\u03081 = \u27e8e1, 1 m B(xop)Stra2pu2p(op)\u27e9, (22) r\u03082 = \u27e8e2, 1 m B(xop)Stra2pu2p(op)\u27e9, (23) where c \u2208 R: a constant (r\u03083 = c), 03 = (0, 0, 0)T, and \u27e8 \u00b7 , \u00b7 \u27e9 denotes scalar product. If c = 0, the operating points: (xop, x\u0307op)T and (rop, r\u0307op)T are in constant altitude flights. Further if r\u03081 = r\u03082 = 0 and c = 0 (i.e. r\u03083 = 0), the operating points are called as equilibrium points: (xe, x\u0307e)T and (re, r\u0307e)T, and are in hovering flights. IV. Problem of multirotor flights in the presence of complete propeller motor failures In the presence of complete propeller motor failures, we consider definitions of multirotor remaining motor speed control signal vectors, and especially theorem of providing motor speed control signals for quadrotor flights. Let Vi \u2282 R2p be the subspace spanned by \u03b5Mi [9], pp. 628\u2013632 such that the motor in the ith rotor is completely failed. Let Pi be the projection onto Vi [9], pp. 628\u2013632. The projection Piu of a vector u \u2208 R2p onto Vi is defined as the vector Piu \u2208 Vi satisfying the orthogonality relation: \u27e8u \u2212 Piu,w\u27e9 = 0 for all w \u2208 Vi, (24) \u27e8Piu, \u03b5Mi\u27e9 = \u27e8u, \u03b5Mi\u27e9. (25) Definition 2: In the case that the motor in the ith rotor is completely failed, the motor speed control signal vector of the remaining motors is determined by ui 2p\u22121 = u2p \u2212 Piu2p = i\u22121\u2211 j=1 \u03c92 M j\u03b5M j + 2p\u2211 j=i+1 \u03c92 M j\u03b5M j, (26) where span{\u03b5M1, \u03b5M2, . . . , \u03b5M2p}: the base vectors of a right 2p-dimensional real vector space R2p. Further Frot(ui 2p\u22121) and Ftra(ui 2p\u22121) are respectively given as Frot(ui 2p\u22121) = Srot2pui 2p\u22121 = Si rot2p\u22121ui 2p\u22121, (27) Ftra(ui 2p\u22121) = Stra2pui 2p\u22121 = Si tra2p\u22121ui 2p\u22121, (28) where Si \u03be2p\u22121 is a matrix such that the ith row of S\u03be2p is deleted, (\u03be = rot, tra). Definition 3: In the case that the motors in the i1th and i2th rotors, (1 \u2264 i1 < i2 \u2264 2p) are completely failed, the motor speed control signal vector of the remaining motors is also determined by ui1,i2 2p\u22122 = u2p \u2212 Pi1 u2p \u2212 Pi2 u2p = i1\u22121\u2211 j=1 \u03c92 M j\u03b5M j + i2\u22121\u2211 j=i1+1 \u03c92 M j\u03b5M j + 2p\u2211 j=i2+1 \u03c92 M j\u03b5M j. (29) Further Frot(ui1,i2 2p\u22122) and Ftra(ui1,i2 2p\u22122) are respectively given as Frot(ui1,i2 2p\u22122) = Srot2pui1,i2 2p\u22122 = Si1,i2 rot2p\u22122ui1,i2 2p\u22122, (30) Ftra(ui1,i2 2p\u22122) = Stra2pui1,i2 2p\u22122 = Si1,i2 tra2p\u22122ui1,i2 2p\u22122, (31) where Si1,i2 \u03be2p\u22122 is a matrix such that the i1th and i2th rows of S\u03be2p are deleted, (\u03be = rot, tra). Theorem 3: In the case of p = 2, let F(x,u2p) be a C1 function taking values in R2p with a neighborhood of x\u0303 in R3 and u\u03032p in R2p and F(x\u0303, u\u03032p) = 02p = (0, 0, . . . , 0)T including both under parts of Eq. (19), and Eq. (20) as follows: F(x, u2p) = A2p(\u03b7)u2p \u2212 b2p, (32) A2p(\u03b7) \u2208 R2p\u00d72p = ( Z(\u03b7)Srot2p eT 3 1 m B(x)Stra2p ) , (33) where c \u2208 R: a constant, b2p = (\u03c8\u0308, \u03b8\u0308, \u03d5\u0308, c + g)T \u2208 R2p. In the case that the motor in the ith rotor is completely failed, we have the following relations: F(x,ui 2p\u22121) = Ai 2p\u22121(\u03b7)ui 2p\u22121 \u2212 b2p\u22121, (34) Ai 2p\u22121(\u03b7) \u2208 R(2p\u22121)\u00d7(2p\u22121) = \u03b5T 2 Z(\u03b7)Si rot2p\u22121 \u03b5T 3 Z(\u03b7)Si rot2p\u22121 eT 3 1 m B(x)Si tra2p\u22121 , (35) b2p\u22121 = (\u03b8\u0308, \u03d5\u0308, c + g)T \u2208 R2p\u22121. (36) In the case that the motors in the i1th and i2th rotors, (1 \u2264 i1 < i2 \u2264 2p) are completely failed, we have the following relations: F(x,ui1,i2 2p\u22122) = Ai1,i2 2p\u22122(\u03b7)ui1,i2 2p\u22122 \u2212 b2p\u22122, (37) Ai1,i2 2p\u22122(\u03b7) \u2208 R(2p\u22122)\u00d7(2p\u22122) = \u03b5T j Z(\u03b7)Si1,i2 rot2p\u22122 eT 3 1 m B(x)Si1,i2 tra2p\u22122 , (38) b2p\u22122 = (\u03b6, c + g)T \u2208 R2p\u22122, (39) where i1 , i2, 2 \u2264 j \u2264 3, j \u2208 N. If j = 2 and j = 3, then \u03b6 = \u03b8\u0308 and \u03b6 = \u03d5\u0308, respectively. If for arbitrary x\u0303 \u2208 R3, det(Ai1,i2 2p\u22122(\u03b7)) , 0, u\u0303i1,i2 2p\u22122 \u2208 R2p\u22122 such that F(x\u0303, u\u0303i1,i2 2p\u22122) = 02p\u22122 is uniquely obtained as u\u0303i1,i2 2p\u22122 = Ai1,i2 2p\u22122(x\u0303)\u22121b2p\u22122. (40) Then \u03c8\u0308, \u03b8\u0308, and \u03d5\u0308 are determined as \u03c8\u0308 = \u27e8\u03b51,Y(\u03b7, x\u0307) + Z(\u03b7)Srot2pu\u0303i1,i2 2p\u22122\u27e9, (41) \u03b8\u0308 = \u27e8\u03b52,Y(\u03b7, x\u0307) + Z(\u03b7)Srot2pu\u0303i1,i2 2p\u22122\u27e9, (42) \u03d5\u0308 = \u27e8\u03b53,Y(\u03b7, x\u0307) + Z(\u03b7)Srot2pu\u0303i1,i2 2p\u22122\u27e9. (43) Proof 1: In the case of p = 2, when we set \u03b7\u0303 \u2208 R2 as a fixed value vector, A4(\u03b7\u0303) becomes a 4 \u00d7 4 constant matrix. Further, in the case of det(Ai1,i2 2p\u22122(\u03b7)) , 0, u\u0303i1,i2 2p\u22122 is clearly obtained. At last by using the second row equation of Eq. (17), \u03c8\u0308 \u03b8\u0308, and \u03d5\u0308 are given in a straightforward way. Thus Theorem 3 is proved. V. Flight simulations of a quadrotor vehicle experiencing motor failures In this section, we present typical flight simulations of a quadrotor vehicle experiencing motor failures as shown in Fig. 6\u20139. \u03b8(t) and \u03d5(t) in Fig. 6 and Fig. 8 are completely overlapped. r1(t) and r2(t) in Fig. 7 and Fig. 9 are also completely overlapped. The following results are obtained by Theorem 1\u2013 3 on MATLAB numerical simulations with an ode45 solver. The needed parameters for numerical computations of the quadrotor are shown in Fig. 5. Under the simulation condition in case that the motor in the first rotor is completely failed: \u03b8 = \u03d5 = 0, kF2 = kF3 = kF4 = 1.79 \u00d7 10(\u22127), kM2 = kM3 = kM4 = 4.38 \u00d7 10(\u22129), c = \u22120.24, we have the following flight simulation (\u03c8\u0308 = 12.03512 as the result) achieved by the motor speed control signals of \u03c9M2 = \u03c9M4 = 6652.2505 [rpm], \u03c9M3 = 0 [rpm] as shown in Fig. 6 and Fig. 7. Under the simulation condition in the case that the motors in the first and third rotors are completely failed: \u03b8 = 0, \u03d5 = 0, kF2 = kF4 = 1.79 \u00d7 10(\u22127), kM2 = kM4 = 4.38 \u00d7 10(\u22129), c = 0, we have the following flight simulation (\u03c8\u0308 = 12.3371 as the result) achieved by the motor speed control signals of \u03c9M2 = \u03c9M4 = 6735.1766 [rpm] as shown in Fig. 8 and Fig. 9. VI. Conclusion The following results are obtained: (1) We have provided definitions of multirotor motor speed control signal vectors, and theorem of directly providing motor speed control signals for quadrotor flights, in the presence of complete motor failures. (2) We have illustrated the typical flight simulations of a quadrotor vehicle experiencing motor failures by Theorem 3. (3) In case that either two motors, or one motor is completely failed, we have confirmed that the determination of yaw angle: \u03c8 must be abandoned in order to keep the altitude. (4) From the flight simulations with the motor failures, we have confirmed that the controls of the altitude and the translational motion are achieved by using two remaining motors, and the control accuracies are less than those by using four motors. (5) We will develop Theorem 3 for a hexarotor, or an octorotor flight in the presence of complete propeller motor failures. References [1] M.W. Mueller and R. D\u2019Andrea, \u201cStability and control of a quadrocopter despite the complete loss of one, two, or three propellers,\u201d 2014 IEEE International Conference on Robotics & Automation (ICRA) Hong Kong Convention and Exhibition Center May 31 \u2013 June 7, Hong Kong, China, pp. 45\u201352, 2014. [2] A. Freddi, A. Lanzon, and S. Longhi, \u201cA feedback linearization approach to fault tolerance in quadrotor vehicles,\u201d 18th IFAC World Congress, pp. 5413\u20135418, Milano Italy August 28 \u2013 September 2, 2011. [3] S. Dongjie, Y. Binxian, and Q. Quan, \u201cReliability Analysis of Multicopter Configurations Based on Controllability Theory,\u201d Proceedings of the 35th Chinese Control Conference July 27-29 Chengdu, China, pp. 6740\u20136745, 2016. [4] M. Saied, B. Lussier, I. Fantoni, H. Shraim, and C. Francis, \u201cFault Diagnosis and Fault-Tolerant Control of an Octorotor UAV using motors speeds measurements,\u201d IFAC PapersOnline 50-1(2017) 5263-5268. [5] V.I. Arnold, Book Title in Mathematical Methods of Classical Mechanics, Second Edition. Springer-Verlag, 1997. [6] Y. Yamamoto, Book Title in From Vector Spaces to Function Spaces - Introduction to Functional Analysis with Applications -, pp. 203\u2013207, SIAM, 2012. [7] H. Okazaki, K. Isogai, and H. Nakano, \u201cModeling and Simulation of Motion of a Quadcopter in a Light Wind,\u201d 2016 IEEE 59th International Midwest Symposium on Circuits and Systems (The IEEE MWSCAS 2016), pp. 117\u2013120, 16\u201319 October 2016, Abu Dhabi, UAE. [8] K. Isogai, H. Nakano, and H. Okazaki, \u201cModeling and simulation of motion of a multirotor in a light wind,\u201d IEICE Tech. Rep., vol. 117, no. 300, CAS2017\u201341, pp. 3\u20138, Nov. 2017 in Japanese. [9] E. Eriksson, D. Estep, and C. Johnson, Book Title in Applied Mathematics: Body and Soul, vol. 2, pp. 621\u2013626, Springer Verlag, 2004." + ] + }, + { + "image_filename": "designv11_80_0000670_pierm19040405-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000670_pierm19040405-Figure1-1.png", + "caption": "Figure 1. Geometry of an 8-pole/12-slot SMPM motor.", + "texts": [ + " In order to simplify the problem, the following assumptions are made: - End effects are neglected. - Stator and rotor iron cores are infinitely permeable. - Radially magnetized magnets with a relative recoil permeability. - The stator slots have radial sides. According to the structural characteristics of the motor and material property, the whole solution domain is divided into three parts: air-gap domain I, permanent magnet domain II, and slot domain Si (i = 1, 2, 3...Q). The motor physic model and its solution domain distribution in the 2-D polar coordinate system are shown in Fig. 1. In Fig. 1, the geometrical parameters are the outer radius of the rotor yoke Rr, the outer radius of the permanent magnet Rm, the inner radius of the stator yoke Rs, and the outer radius of the slot-opening Rsy. The slot-opening angle is \u03b2, and \u03b8i is the mechanical position of the rotor. In the 2-D polar coordinate, only Z axis has the virtual value when PM is radial magnetization. Its magnetization expression is r and \u03b8 function which is derived by the Fourier decomposition [1], M(r, \u03b8) = \u23a1 \u23a3 \u221e\u2211 v=1,3,5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001182_s13344-019-0054-0-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001182_s13344-019-0054-0-Figure1-1.png", + "caption": "Fig. 1.\u00a0\u00a0\u00a0AUV\u00a0layout\u00a0and\u00a0the\u00a0reference\u00a0frames.", + "texts": [ + "\u00a0The\u00a0AUV\u00a0that\u00a0is\u00a0studied\u00a0in the\u00a0paper\u00a0is\u00a0also\u00a0introduced\u00a0in\u00a0this\u00a0section;\u00a0in\u00a0Section\u00a03\u00a0the synthesis\u00a0and\u00a0analysis\u00a0of\u00a0the\u00a0depth\u00a0controller\u00a0explained; Section\u00a04\u00a0presents\u00a0the\u00a0proposed\u00a0identification\u00a0method;\u00a0in Section\u00a05\u00a0the\u00a0sensor\u00a0fusion\u00a0algorithm\u00a0is\u00a0developed;\u00a0in\u00a0Section\u00a06\u00a0a\u00a0six-degree-of-freedom\u00a0simulation\u00a0of\u00a0AUV\u00a0is\u00a0performed\u00a0and\u00a0the\u00a0performance\u00a0of\u00a0identification\u00a0and\u00a0fusion\u00a0algorithms\u00a0are\u00a0validated;\u00a0the\u00a0last\u00a0section\u00a0concludes\u00a0the\u00a0paper. 2 Dynamic modelling The\u00a0fusion\u00a0and\u00a0identification\u00a0algorithms\u00a0are\u00a0going\u00a0to\u00a0be developed\u00a0for\u00a0the\u00a0AUV\u00a0shown\u00a0in\u00a0Fig.\u00a01.\u00a0The\u00a0values\u00a0of\u00a0the dynamic\u00a0parameters\u00a0are\u00a0determined\u00a0using\u00a0CFD\u00a0and\u00a0Computer-Aided\u00a0Design\u00a0(CAD)\u00a0methods.\u00a0The\u00a0mass\u00a0of\u00a0the\u00a0AUV is\u00a050\u00a0kg;\u00a0the\u00a0overall\u00a0length\u00a0is\u00a01.94\u00a0m;\u00a0its\u00a0maximum\u00a0diameter\u00a0is\u00a017.6\u00a0cm.\u00a0The\u00a0vehicle\u00a0has\u00a0four\u00a0control\u00a0fins\u00a0of\u00a0type NACA0009.\u00a0The\u00a0AUV\u00a0displaces\u00a00.0493\u00a0m3\u00a0which\u00a0means that\u00a0it\u00a0has\u00a0neutral\u00a0buoyancy\u00a0in\u00a0sea\u00a0water\u00a0with\u00a0a\u00a0density\u00a0of 1015\u00a0kg/m3.\u00a0The\u00a0center\u00a0of\u00a0gravity\u00a0(CG)\u00a0of\u00a0the\u00a0vehicle\u00a0is\u00a0exactly\u00a01\u00a0cm\u00a0beneath\u00a0the\u00a0center\u00a0of\u00a0buoyancy. The\u00a0equations\u00a0of\u00a0motion\u00a0of\u00a0the\u00a0AUV\u00a0are\u00a0coupled\u00a0and nonlinear\u00a0(Ernani\u00a0et\u00a0al.,\u00a02015).\u00a0Six\u00a0degrees\u00a0of\u00a0freedom\u00a0(6DOF)\u00a0differential\u00a0equations\u00a0can\u00a0be\u00a0obtained\u00a0using\u00a0Newton and\u00a0Euler\u00a0equations\u00a0(Fossen,\u00a01994).\u00a0The\u00a0equations\u00a0are\u00a0derived\u00a0employing\u00a0two\u00a0coordinate\u00a0frames\u00a0(Fig.\u00a01):\u00a0the\u00a0body frame\u00a0whose\u00a0origin\u00a0coincides\u00a0with\u00a0the\u00a0center\u00a0of\u00a0buoyancy, and\u00a0the\u00a0earth\u00a0fixed\u00a0frame\u00a0(Hong\u00a0et\u00a0al.,\u00a02013). u v w p q r X Y Z K M N \u03c6 \u03b8 \u03c8 The\u00a0linear\u00a0velocity\u00a0components\u00a0[ ,\u00a0 ,\u00a0 ]\u00a0(surge,\u00a0sway, heave),\u00a0the\u00a0angular\u00a0velocities\u00a0[ ,\u00a0 ,\u00a0 ]\u00a0(roll,\u00a0pitch,\u00a0yaw rates),\u00a0the\u00a0external\u00a0forces\u00a0[ ,\u00a0 ,\u00a0 ],\u00a0and\u00a0the\u00a0external\u00a0moments\u00a0[ ,\u00a0 ,\u00a0 ]\u00a0are\u00a0defined\u00a0in\u00a0the\u00a0body\u00a0frame.\u00a0On\u00a0the\u00a0other\u00a0hand,\u00a0transitional\u00a0displacements\u00a0[x,\u00a0y,\u00a0z]\u00a0and\u00a0attitude Euler\u00a0angles\u00a0[ ,\u00a0 ,\u00a0 ]\u00a0are\u00a0defined\u00a0using\u00a0the\u00a0earth\u00a0fixed frame\u00a0(Kim\u00a0et\u00a0al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure2-1.png", + "caption": "Figure 2. The modeling of impeller.", + "texts": [ + " Using CAE analysis software to illustrating the results of top gating system, bottom gating system and two side gating systems. Furthermore, with the assistance of CAE analysis, the gating system will be optimized and the best project of CAE simulation will be practiced. Before the practice of manufacture, AnyCasting is applied for design of gating system. The parameters of process are chosen depend on reality fabricate condition of workshop. Thus the results of production will only due to the management of different gating system. The 3D model is generated according to the setting of size in figure 2. Then the model that including four types of gating system are created. Top gating system, bottom gating system and 2 types of side gating system are added to the impeller as Figure 3. ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 2.2.Simulation of gating system The software Anycasting bears the simulation. The mesh are created in the preprocessing module called AnyPRE. Taking the top gating system as the example, the account of flexible finite elements is 103488 in Figure 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure9-1.png", + "caption": "Figure 9. The results that solidification of model with different number of risers.", + "texts": [ + " All the filling and solidification are stable, and effect-free defects in gating system and shrinkage defects in risers, the advance of side gating system is obvious. However, the solidification in cross-sharp gating system is more stable. As well as defect range in cross-sharp is smaller than the other. So it is worth for further improvements of cross-sharp gating system. The last solidification in cross-sharp gating system is zone in yellow like Figure 8(a). but this area is far away from the riser. This state indicate that the managements of risers can be performed. The Figure 9 describes the solidifications sequence of mould with more risers. The white zone means the part that solidified lastly. Only in the Figure 9(d). all the white area located in the flow ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 channel, which means the well sequence of solidification. Therefore, the best solution of gating system is cross-sharp side gating system with four risers. 3.2 Experiment and confirmation Mix the thermo stability plaster with water as proportion 100:46 and paint the plaster onto the 3D printed mould. Put it into the casting mold when the plaster is dried" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003461_embc44109.2020.9175945-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003461_embc44109.2020.9175945-Figure3-1.png", + "caption": "Fig. 3. Mechanism of for opening and closing three fingers", + "texts": [ + " The relationship between a force Fs applied to the small-diameter syringe in the operating interface and a force Fl applied to the large-diameter syringe in the gripper to open the fingers is expressed as follows: Fl = Sl Ss \u00d7 Fs, (1) where Sl and Ss is the cross-sectional area of the large- diameter syringe and the small-diameter syringe, respectively. In the developed gripper, the force applied to the smalldiameter syringe in the operating interface is amplified about 2.1 times as follows: Fl = 2.13Fs. (2) Fig. 3 shows the inside of the developed gripper. We used a basic structure of the three fingers developed by our research group in the past [12]. Each finger is connected to the center connector part with one joint. Three fingers are simultaneously opened by pushing the connector with the plunger with a stroke of 10 mm. The polyacetal guide pin of the finger moves in the guide track, which controls the trajectory of the finger. Conversely, three fingers are closed by returning the connector to the initial position by two extension springs (spring constant 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003840_3424978.3425059-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003840_3424978.3425059-Figure1-1.png", + "caption": "Figure 1: Schematic Diagram of Stinger Test System.", + "texts": [ + " Based on the theory of semi physical simulation, this section designs and completes the indoor meso large scale stinger model test, obtains the load and response of each position of the stinger mechanism under different working conditions, and discusses the related testing technology in detail. The overall device of the model test consists of the model of the stinger, the model of the idler box, the model of the pipeline, the vibration table, the sensor, the dSPACE control system, the device for providing the pipeline tension, etc. The schematic diagram of test principle is shown in Figure 1, and Figure 2 is the physical diagram of stinger test. As shown in Figure 1, during the test, the vibration of the pipelaying ship on the sea is actually measured, and the vibration of the sea wave is transformed into the equivalent actuator data, in which the actuator directly below is used to simulate the sinking and floating of the ship, and the actuator on the second floor is used to simulate the pitching and rolling of the ship, so the simulation of the sea situation is completed by using the computer-controlled actuator. Then the force and deformation of the model are measured by measuring equipment, and the load of the stinger on the ocean can be obtained by conversion relationship" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003110_s40194-020-00955-7-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003110_s40194-020-00955-7-Figure7-1.png", + "caption": "Fig. 7 Cloud images of powder transport ratio in (a) \u03c7-\u03b1, (b) \u03c7-D, and (c) \u03b1-D coordinate system", + "texts": [ + " A total of three groups of measurements were developed, and the data of 12 measurement points are listed in Table 3. In Fig. 4(b), a powder collection device equipped with two steel sheets, springs, and hexagon socket bolts was used to verify the powder transport ratio under different powder injection parameters. All tests were programmed with 2.322 g/min powder feed rate during 10 s and the placement was marked to ensure that each test has the same initial state before parameter adjustment. As the influence of molten pool width had been developed and compared with theoretical values in Fig. 7 of [8], powder transport ratio had been measured only after adjusting the defocusing distance and powder feeding angle respectively, which ranged from \u2212 2 to 10 mm by 2 mm increments for \u03c7, and from \u2212 3\u00b0 to 3\u00b0 by 0.5\u00b0increments for \u03b1. By weighing the powder quantity stored between above two steel sheets always set at 3 mm distance, the experimental data could be estimated and distributed in Fig. 6. Similar to the experiment in Section 3.1 before, adjustment of defocusing distance \u03c7 (+ 4 mm) and powder feeding angle \u03b1 (+ 3\u00b0) was substituted into the theoretical model successively", + " This can be both observed from theoretical and experimental results, which means this established model can describe the influence of powder feeding parameters on real powder efficiency. According to Section 2.4, \u03bePT values corresponding to 13 \u00d7 13 \u00d7 5 levels of three variables (\u03c7, \u03b1, \u03b4) were calculated. These three-dimensional point clouds are displayed in the form of some oriented two-dimensional cloud images to reflect the influence of the powder feeding parameters on \u03bePT, which is also illustrated in Fig. 7 for their orientation relationships. As shown in Fig. 7(b, c), \u03bePT increases with increasing molten pool width under any conditions, and the maximum value of this increment, which is related to the defocusing distance or powder feeding angle, occurs near the initial position without change of \u03c7 or \u03b1. As mentioned before, the positive and negative adjustments for the above parameters are asymmetrical. In Fig. 7(a), \u03bePT value can be maintained at a higher (/stable) level by simultaneous adjustments of \u03c7 and \u03b1 in negative (/positive) direction, but a lower level as (\u03c7 > 0, \u03b1 < 0) or (\u03c7 < 0, \u03b1 > 0). For the physical model, as the lateral nozzle is located at the position where central particles can fall into the molten pool center, the powder efficiency increases significantly (/slightly) with the reduction (/increase) of defocusing distance and powder feeding angle simultaneously. From the above analysis, a qualitative judgment on the variation trend of \u03bePT can be made by the direction (+, \u2212) of the adjusted parameters (\u03c7, \u03b1, \u03b4) preliminarily" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000042_la-cci.2018.8625204-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000042_la-cci.2018.8625204-Figure1-1.png", + "caption": "Fig. 1: Quadrotor diagram.", + "texts": [], + "surrounding_texts": [ + "The model of the quadrotor is derived using the EulerLagrange methodology. Nonlinear dynamics of the quadrotor are described in terms of applied forces and moments." + ] + }, + { + "image_filename": "designv11_80_0000948_ever.2019.8813582-Figure17-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000948_ever.2019.8813582-Figure17-1.png", + "caption": "Fig. 17. Illustration of unequal rotor teeth.", + "texts": [ + " As aforementioned, the open-circuit DC winding induced voltage only contains the 6th harmonic and its multiples in the 12-slot/10-pole HESFPM machine, in which the 6th harmonic is the dominant harmonic. Both the amplitude and initial phase of the 6th harmonic varies with rotor outer pole arc as shown in Figs. 15 and 16. As a consequence, it is expected that the open-circuit DC winding induced voltage can be reduced by unequal rotor teeth. Inspired by this phenomenon, unequal rotor teeth is employed to reduce the open-circuit DC winding induced voltage as illustrated in Fig. 17. The mechanism behind this technique is to cancel the corresponding induced voltage harmonics by a pair of rotor outer pole arcs. All available combinations of two rotor pole arcs are calculated by FE method. The peak to peak values of the open-circuit DC winding induced voltages and the armature winding phase fundamental back-EMFs are plotted against two rotor outer pole arcs as shown in Figs. 18 and 19. In order to reduce the open-circuit DC winding induced voltage as much as possible, whilst keep the average electromagnetic torque as higher as possible, two selection criteria are adopted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001552_s12555-019-0234-y-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001552_s12555-019-0234-y-Figure2-1.png", + "caption": "Fig. 2. (blue) Rolling- and (red) pin-based discrete bending joints compared to the (black) ideal continuum joint.", + "texts": [ + " Section 2 explains the behavior of a 1-DOF bending joint. Next, the advantages of the proposed 2N+1 method for 2-DOF bending are demonstrated quantitatively compared to the existing 2N methods in Section 3. Section 4 presents a performance review of the proposed bending joint by analyzing an example of suturing motion. Finally, the utility of the proposed joint mechanism is discussed in Section 5. 2. 1-DOF DISCRETE BENDING JOINT Before discussing the 2-DOF isotropic bending joint, it is necessary to understand the behavior of the 1-DOF joint. Fig. 2 shows the fine positioning differences among the distal tips of three types of bending joints at the same bending angle: continuum bending, and pin- and rollingbased discrete bending joints. Assuming that the lengths of the central lines remain constant, between the rollingbased discrete bending joint and the pin-based discrete bending joint, the central line path and distal tip position of the rolling-based discrete bending joint are closer to those of the ideal continuum joint. Because each rolling unit is a combination of two adjacent identical pin joints, rolling-based bending results in a finer bending motion than pin-based bending with the same number of units [24]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003112_j.mechmachtheory.2020.103903-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003112_j.mechmachtheory.2020.103903-Figure8-1.png", + "caption": "Fig. 8. Screenshot, from digital simulation in video file Xactuator.mp4 in supplementary materials, of 5-DOF 2-coupled-Cartesian-manipulator with nonintersecting-revolute-axes and non-parallel-revolute-axes.", + "texts": [ + " (12 , 13 ) A n W = T W \u2212 R W B n ZXY T Bn \u2212 R W A n ZX B n An (18) C n W = T W \u2212 R W D n ZXY T Dn \u2212 R W C n ZX D n Cn (19) Given desired common-link L T position T W and orientation \u02c6 Z W T , all of the terms on the right hand sides of Eqs. (18 , 19 ) are known from equations (5 , 6 , 14-17 ) and the fixed geometry of the links B n An , T Bn , D n Cn , T Dn . Video file, Xactuator.mp4 in supplementary materials, illustrates a digital simulation example of the inverse kinematics Eqs. (18 , 19 ). Fig. 8 is a screenshot from the simulation. Inverse kinematics of 2-coupled-Cartesian-manipulator with intersecting-revolute-axes. For the intersecting-revolute-axes case, B n An = [ 0 0 0 ] \u2032 , D n Cn = [ 0 0 0 ] \u2032 , T Bn = [ 0 0 z Bn T ] \u2032 , T Dn = [ 0 0 z Dn T ] \u2032 . This simplifies the inverse kinematics, Eqs. (18 , 19 ) to, A n W = T W \u2212 z Bn T \u02c6 Z W T (20) C n W = T W \u2212 z Dn T \u02c6 Z W T (21) Conceptually Eqs. (20 , 21 ) correspond to controlling the position T W and orientation \u02c6 Z W of a line in 3D space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001312_0954407019877302-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001312_0954407019877302-Figure7-1.png", + "caption": "Figure 7. (a and b) Front and (c and d) rear view of Q-criterion (4000 s 2) isosurface with velocity line integral convolution vectors and turbulent charge contour of the solid wheel from time-averaged URANS simulation. Jetting phenomena, recirculation bubble and rim vortices evolution in time are captured. Surprising transversal vortex on the rear contact is found: (a) u = 0 , (b) u = 360 , (c) u = 0 , and (d) u = 360 .", + "texts": [ + " The pressure plateaus between u=1008 and u=2208 corresponding to a region are characterised by high turbulent intensity; CFD is in overall good agreement with measurement. The most challenging feature to capture is the separation location and strength. The pressure dips at u\u20192808, suggesting the start of the separation. CFD simulations predict sooner and stronger separation compared with experiment; this trend is also reported in the literature, by Mears and Dominy15 at u=2908 and closer to u=2708 by Knowles,11 Hinson,12 Heyder-Bruckner17 and McManus and Zhang.10 Near-wake behaviour is shown in Figure 7; Q-criterion isosurfaces flooded by turbulent charge field function n=r L where lambda vector is defined as the cross product of velocity and vorticity L[v3r3v. Two time strands are shown as follows: at initial state of one revolution and its final state, in frontal and back isometric views. The jetting effect and rim vortices can be observed on both time intervals; the jet is periodically turned into arch-shaped vortices that propagate downstream. The recirculation zone predicted by the pressure coefficient can also be observed in the rear isometric view of Figure 7, with the recirculation bubble in the near wake. The incoming flow leaves the top of the tyre, from both sides of the rim, and is merged into an arch-shaped vortex in the rear. A long and thin vortical structure is observed downstream of the contact, spanning along the width of the wheel, as shown in Figure 7(c). This might suggest a starting vortex structure formation. To the best of the authors\u2019 knowledge, this feature has not been observed in previous studies and is one of the main differences with McManus and Zhang10 results. This could be simply be a numerical artefact; however, models with elevated contact patch are not able to capture this flow feature. Nonetheless, turbulent scale resolving methods such DES could enhance our understanding on this as experimental data points are extremely difficult to obtain near this region" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure8-1.png", + "caption": "Figure 8. Pad boundary conditions", + "texts": [], + "surrounding_texts": [ + "The\ufeff pad\ufeff is\ufeff decomposed\ufeff into\ufeff 4\ufeff nodes\ufeff tetrahedral\ufeff elements\ufeff due\ufeff to\ufeff the\ufeff existence\ufeff of\ufeff a\ufeff groove\ufeffwith\ufefflines\ufeffand\ufeffarches\ufeffas\ufeffwell\ufeffas\ufefforifices\ufeffin\ufeffthe\ufeffcircle\ufeffform.\ufeffThese\ufeffshapes\ufeffrequire\ufefftetra\ufeff elements\ufeffto\ufeffachieve\ufeffthe\ufeffmost\ufeffaccurate\ufeffgeometry\ufeffoverlap,\ufeffFigure\ufeff6.\ufeffThe\ufeffbad\ufeffis\ufeffdecomposed\ufeffat\ufeff 12\ufeffpoints\ufeffin\ufeffthe\ufeffaxial\ufeffdirection,\ufeffthe\ufeffangular\ufeffamplitude\ufeffhas\ufeff70\ufeffpoints\ufeffand\ufeff4\ufeffnodes\ufeffdepending\ufeff on\ufeff the\ufeff thickness.\ufeff The\ufeff feed\ufeff groove\ufeff is\ufeff discretized\ufeff at\ufeff 13\ufeff points\ufeff in\ufeff the\ufeff radial\ufeff direction\ufeff and\ufeff 5\ufeffnodes\ufeffalong\ufeff its\ufeffwidth.\ufeffThis\ufeff finite\ufeffelement\ufeffmesh\ufeffof\ufeff the\ufeffplain\ufeffbearing\ufeffconsists\ufeffof\ufeff19830\ufeff elements\ufeffand\ufeff47784\ufeffnodes.\nBoundary Conditions To\ufeffbetter\ufeffresolve\ufeffthe\ufeffequations\ufeffpresented\ufeffearlier\ufeffin\ufeffthis\ufeffarticle,\ufeffit\ufeffis\ufeffnecessary\ufeffto\ufeffapply\ufeffappropriate\ufeff boundary\ufeffconditions.\ufeffThe\ufeffboundary\ufeffconditions\ufeffapplied\ufeffto\ufeffthe\ufeffshaft\ufeffare:\n\u2022\ufeff The\ufeffnodes\ufeffof\ufeffthe\ufeffface\ufeffwhere\ufeffthe\ufeffshaft\ufeffis\ufeffcoupled\ufeffwith\ufeffa\ufeffrotating\ufeffmanifold\ufeffare\ufeffblocked\ufeffalong\ufeffthe\ufeff x-axis,\ufeffFigure\ufeff7.\nThe\ufeffboundary\ufeffconditions\ufeffapplied\ufeffto\ufeffthe\ufeffpad\ufeffare:\n\u2022\ufeff The\ufeffpad\ufeffis\ufeffplaced\ufeffin\ufeffa\ufeffsupport\ufeffring\ufeffwhose\ufefflower\ufeffpart\ufeffis\ufeffspherical,\ufeffa\ufeffframe\ufeffone\ufeffbears\ufeffon\ufeffthe\ufeffface\ufeff of\ufeffthe\ufeffspherical\ufeffhydrostatic\ufeffplain\ufeffbearing\ufeffand\ufeffwhich\ufeffis\ufeffembedded\ufeffin\ufeffthe\ufeffbase,\ufeffFigure\ufeff8; \u2022\ufeff The\ufeffpad\ufeffis\ufefflocked\ufeffon\ufeff60\u00b0\ufeffof\ufeffthe\ufefflower\ufeffpart:\ufeffBlocking\ufeffof\ufeffthe\ufeffnodes\ufeffaccording\ufeffto\ufeffx,\ufeffy\ufeffand\ufeffz; \u2022\ufeff The\ufeffpad\ufeffis\ufeffmounted\ufeffin\ufeffa\ufeffring\ufeffso\ufeffit\ufeffis\ufeffcylindrical\ufeffsupport\ufeffwith:\n\ufeff\u25e6 Fixed\ufeffradial\ufeffnodes; \ufeff\u25e6 Free\ufeffAxial\ufeffNodes; \ufeff\u25e6 Fixed\ufefftangential\ufeffnodes.\nInsertion of Pressures We\ufeffcan\ufeffapply\ufeffglobal\ufeffloads,\ufeffstructural,\ufeffas\ufeffwell\ufeffas\ufeffimposed\ufeffdisplacements\ufeffaccording\ufeffto\ufeffthe\ufeffcases\ufeffstudied.\ufeff In\ufeffthe\ufeffcase\ufeffstudied,\ufeffpressures\ufeffare\ufeffapplied\ufeffalong\ufeffthe\ufeffcircumferential\ufeffaxis\ufeffas\ufeffwell\ufeffas\ufeffalong\ufeffthe\ufeffaxial\ufeffaxis\ufeff of\ufeffthe\ufeffshaft\ufeffand\ufeffthe\ufeffplain\ufeffbearing\ufeff(Figure\ufeff9).", + "Radial Load Effect Pressure Distribution for Textured Plain Bearing To\ufeff demonstrate\ufeff the\ufeff effect\ufeff of\ufeff the\ufeff radial\ufeff load\ufeff on\ufeff the\ufeff operating\ufeff performance\ufeff of\ufeff the\ufeff non-textured\ufeff hydrodynamic\ufeffplain\ufeffbearing,\ufeffsuch\ufeffas\ufeffthe\ufeffpressure\ufeffdistribution\ufeffand\ufeffthe\ufefffluid\ufeffflow\ufeffvelocity\ufeffwithin\ufeffthe\ufeff plain\ufeffbearing,\ufeffa\ufeffradial\ufeffload\ufeffvariation\ufeffis\ufeffperformed\ufeff(W1\ufeff=\ufeff2000N,\ufeffW2\ufeff=\ufeff6000N\ufeffand\ufeffW3\ufeff=\ufeff10000N).\ufeff The\ufeffinitial\ufeffoperating\ufeffconditions\ufeffof\ufeffthe\ufeffbearing\ufeffare\ufeffwith\ufeffa\ufeffsupply\ufefftemperature\ufeffTa\ufeff=\ufeff40\ufeff\u00b0C,\ufeffsupply\ufeff pressure\ufeffPa\ufeff=\ufeff0.04\ufeffMPa\ufeffand\ufeffthe\ufeffspeed\ufeffof\ufeffrotation\ufeffof\ufeffthe\ufeffshaft\ufeffequal\ufeffto\ufeffN\ufeff=\ufeff6000\ufeffrpm.", + "Figure\ufeff10\ufeffshows\ufeffthe\ufeffdistribution\ufeffof\ufeffthe\ufeffpressure\ufeffalong\ufeffthe\ufeffmedian\ufeffplane\ufeffof\ufeffthe\ufeffplain\ufeffbearing\ufefffor\ufeff different\ufeffradial\ufeffloads.\ufeffThe\ufeffsignificant\ufeffpressures\ufeffare\ufeffobtained\ufefffor\ufeffa\ufefftextured\ufeffplain\ufeffbearing\ufeffunder\ufeffa\ufeffradial\ufeff load\ufeffof\ufeff10000N.\ufeffAs\ufeffwell\ufeffas\ufeffthe\ufeffincrease\ufeffof\ufeffthe\ufeffradial\ufeffload\ufeffleads\ufeffto\ufeffan\ufeffincrease\ufeffof\ufeffthe\ufeffhydrodynamic\ufeff pressure.\ufeffThis\ufeffincrease\ufeffreaches\ufeff75\ufeffpercent\ufefffor\ufeffthe\ufefftextured\ufeffplain\ufeffbearing\ufeffcase.\nThe\ufeffpressure\ufeffdistribution\ufeffaccording\ufeffto\ufeffangular\ufeffposition\ufefffor\ufeffplain\ufeffbearing\ufefftextured\ufeffand\ufeffnot\ufefftextured\ufeff for\ufeffradial\ufeffload\ufeffof\ufeff10\ufeff000N\ufeffand\ufeffrotational\ufeffvelocity\ufeffof\ufeff9000rpm\ufeffis\ufeffpresented\ufeffrespectively\ufeffin\ufeffFigure\ufeff 10a\ufeffand\ufeffFigure\ufeff11.\ufeffThe\ufeffcurve\ufeffclearly\ufeffshows\ufeffthat\ufeffthe\ufeffpressure\ufeffdistribution\ufeffalong\ufeffthe\ufeffmedian\ufeffplane\ufeffof\ufeff the\ufeffplain\ufeffbearing\ufeffis\ufeffdifferent\ufefffor\ufeffthe\ufeffcase\ufeffof\ufeffa\ufeffplain\ufeffbearing\ufeffwithout\ufefftexture\ufeffand\ufeffa\ufeffplain\ufeffbearing\ufeffwith\ufeff textured\ufeffsurface.\ufeffthe\ufeffdifference\ufeffis\ufeffestimated\ufeff75\ufeffpercent.\nDisplacement Distribution for Textured Bearing The\ufeffelastic\ufeffdisplacement\ufeffof\ufeffthe\ufeffbearing\ufeffinner\ufeffface\ufeffis\ufeffdue\ufeffto\ufeffthe\ufeffdeformation\ufeffof\ufeffthe\ufeffbearing\ufeffgenerated\ufeff by\ufeffthe\ufeffhydrodynamic\ufeffpressure\ufefffield,\ufeffis\ufeffshown\ufeffin\ufeffFigure\ufeff12\ufefffor\ufeffa\ufeffload\ufeff2000\ufeffN\ufeffand\ufeffthe\ufeffbearing\ufeffoperates\ufeff at\ufeffa\ufeffrotational\ufeffvelocity\ufeffof\ufeff6000\ufeffrpm.\ufeffThe\ufeffmaximum\ufeffdeformation\ufeffis\ufeffnoted\ufeffat\ufeffthe\ufefflevel\ufeffof\ufeffthe\ufefflower\ufeff generatrix\ufeffof\ufeffthe\ufeffbushing.\nFigure\ufeff13\ufeffillustrates\ufeffthe\ufeffdisplacement\ufeffdistribution\ufeffin\ufeffthe\ufeffcircumferential\ufeffdirection\ufeffof\ufeffthe\ufeffplain\ufeff bearing\ufeffas\ufeffa\ufefffunction\ufeffof\ufeffthe\ufeffradial\ufeffload\ufeffand\ufefffor\ufeffthe\ufeffrotation\ufeffvelocity\ufeffof\ufeff6000\ufeffrpm.\ufeffThe\ufeffcurve\ufeffshows\ufeff that\ufeffthe\ufeffdeformation\ufeffhas\ufeffa\ufefflarge\ufeffincrease\ufeffwith\ufeffthe\ufeffrise\ufeffof\ufeffthe\ufeffload,\ufeffthis\ufeffincrease\ufeffreaches\ufeff0.35\u03bcm\ufefffor\ufeff a\ufefftextured\ufeffsurface\ufeffplain\ufeffbearing;\ufeffon\ufeffthe\ufeffother\ufeffhand,\ufeffreaches\ufeffthem\ufeff0,16\u00b5m\ufefffor\ufeffa\ufeffnon-textured\ufeffplain\ufeff bearing,\ufefffor\ufeffthe\ufeffcase\ufeffof\ufeffa\ufeffplain\ufeffbearing\ufeffsubmitted\ufeffa\ufeffloading\ufeff10000N.\nFigure\ufeff14\ufeffshows\ufeffthe\ufeffdisplacement\ufeffalong\ufeffthe\ufeffcircumference\ufeffdirection\ufeffof\ufeffthe\ufeffpad.\ufeffDisplacement\ufeff is\ufeffvery\ufeffimportant\ufefffor\ufeffthe\ufeffcase\ufeffof\ufeffa\ufefftextured\ufeffbearing\ufeffand\ufeffvery\ufeffloaded,\ufeffit\ufeffforms\ufeffa\ufeffkind\ufeffof\ufeffa\ufeffsignificant\ufeff hollow\ufeffat\ufeffthe\ufefflevel\ufeffof\ufeffits\ufefflower\ufeffbearing\ufeffgeneratrix\ufeffcompared\ufeffto\ufeffthat\ufeffnoted\ufefffor\ufeffa\ufeffnon-textured\ufeffbearing.\ufeff" + ] + }, + { + "image_filename": "designv11_80_0003148_012030-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003148_012030-Figure8-1.png", + "caption": "Figure 8. AGV on the line", + "texts": [ + " From Figure 7, the AGV is set to follow one of the following paths: From A to B From A to C From B to A From C to A Note that there is no path set for the AGV to go form B to C or from C to B because each point represents the location of components of the entire system. Point A: is the location of the storeroom Point B: is the location of the first Assembly line Point C: is the location of the second Assembly line With that in mind, there is no need for the AGV to travel from one assembly line to the other. The decision process of the AGV mostly depends on the inputs of the sensors that are used to follow the line. The commands form the masters just tells the AGV where it must go. Observe Figure 8, the AGV is placed on the line with 2 sensors on the left of the line and 2 sensors on the right of the line. 2nd 2019 ICERA Journal of Physics: Conference Series 1577 (2020) 012030 IOP Publishing doi:10.1088/1742-6596/1577/1/012030 To follow the line, the AGV makes use of only 2 sensors (R1 and L1). These sensors have a default value of 1 and changes to 0 when the sensor passes over the black line. The other 2 sensors are used as check point counter. The check points next to the black line are utilized by the AGV to determine its location compared to where it is going" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001090_032034-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001090_032034-Figure6-1.png", + "caption": "Figure 6. Diagram of satellite measurements (a) on the outer and dividing cone, (b) the inner hole.", + "texts": [ + " The setting up of tools is carried out using special standards and standard details available at workplaces. To improve the effectiveness of the methodology [7], as well as the possibility of using the correlation method, a measurement scheme of the studied part was developed. Measurements of the billet and the processed semi-finished product are performed in a single coordinate system, for which its beginning is marked on the billet as a mark of the first tooth. Measurement of the beats on the outer and dividing cone is performed for each tooth cavity, which receives its own sequence number (fig. 6 a). Measurements of the internal diametral size of the pre-bar element of the billet are made in two mutually perpendicular sections along the length along the length of the billet. The location of these sections is also known by reference to the numbers of the troughs (fig. 6 b). Mechanical Science and Technology Update IOP Conf. Series: Journal of Physics: Conf. Series 1260 (2019) 032034 IOP Publishing doi:10.1088/1742-6596/1260/3/032034 The research method consists in the consistent application of various methods of analysis: 1. The primary analysis of measurement data. At this stage, the measurement data is checked for gross errors according to the Grubbs criterion. The histogram is executed. Preliminary check of the current law of distribution and calculation of statistical characteristics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure9-1.png", + "caption": "Figure 9. Stephenson2 mechanism (1).", + "texts": [], + "surrounding_texts": [ + "The primary objective of this study was to identify the joint that producesthe highest error in a mechanism due to clearance. Simulations were carried out for all mechanisms described in the previous section by introducing clearances in different joints, one at a time. The effect of different speeds (of crank) and clearance sizes was also studied. Some important results are discussed here. The results for the four bar mechanisms have been summarized in figures 13, 14 and 15. These show that Joint 3 is the most critical joint and the values vary significantly for 0.02 mm but they are very similar for 0.5 mm. So, sensitivity to speed is high for low clearance. Figures 16, 17 and 18 show mean deviation at different joints of different inversions of the Watt chain at 0.1 mm radial clearance. The trend of all the curves is the same. Joint 3 is the most critical joint in all the three mechanisms followed by joints 2, 4 and 5, i.e., the criticality order here is: Joint 3[ Joint 2[Joint 4[ Joint 5. Joint 6 in figures 5 and 6 is the grounded joint which is least critical of them all. Figures 19 and 20 show the criticality graphs for Stephenson1 mechanism and Stephenson1 mechanism with slider, respectively. Both the graphs show the same trend where the deviation is maximum at the output link and decreases as one proceeds to the other end of the mechanism. A common trend which seems to emerge is that the first joint next to the crank, which has a link connected to ground, is the most critical joint. After that the criticality ranking goes towards the crank and then goes further down to the joints in order of their distances from the crank. We would examine the validity of this hypothesis in the rest of the study. Figures 21 and 22 show the criticality graphs for Stephenson2 mechanism (1) and Stephenson2 mechanism (2). The order of criticality for both the mechanisms is: Joint4 [ Joint3 [ Joint2 [ Joint5. This is as per the hypothesis stated earlier. The most critical joint is the joint next to the crank, which has a link connected to the ground (i.e., Joint 4). The next most critical joints are the ones towards the crank, i.e., Joint 3 and Joint 2. Then comes the joints which are left, i.e., Joint 5. Joint 6 is the grounded joint. The two graphs look very similar as the structures of mechanisms are the same. The deviations increase significantly with increase in speed, if the clearance is at Joint 4. The values in first graph are slightly higher than the corresponding values in the second graph. The overall conclusion is that deviations are more, if the output link is closer to the crank. Similar results are seen in eight and ten-bar mechanisms as shown in figures 23 and 24. The joint next to the crank is the most critical joint. It should also be noted that the grounded joints also show a trend. The grounded joint in the first loop has a significant effect but rest of the grounded joints have very little effect on the output. Error in output generally increases with increase in the crank speed (expect for an aberration shown in figure 19). This happens for all clearance sizes. However, the change in mean deviation is quite small even with a 5 times increase in speed from 100 to 500 rpm. Figures 25 and 26 show the variation of mean deviation with increase in speed when a clearance of 0.1 mm is at Joint 2 and Joint 3, respectively. It can be noticed that the deviations usually increase with speed. This means that a machine working at higher speeds is expected to produce more error than a machine working at lower speeds. Increase in radial clearance size brings an almost linear increase in the deviation values as seen in figures 27 and 28. It is also worth noting from figure 27 here that the deviation of slider is nearly 50% for 0.1 mm clearance (0.0501 is 50% of 0.1) and it goes on increasing to 80% for 0.5 mm clearance (0.3977 is 80% of 0.5). This happens because at lower values of clearance, the number of collisions between journal and bearing is more. Due to more number of collisions, the journal is sent back to the center again and again, and thus it remains at the center for a larger amount of time as compared to the case of large clearances. Thus, in larger clearances, the journal appears to remain closer to the walls of the bearing for an extended period and spends less time going to the diametrically opposite wall." + ] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.1-1.png", + "caption": "Figure 1.1. Rotation of angle ! in a plane", + "texts": [ + " For three-dimensional modeling, it is essential to have a good understanding of the concepts related to rotation matrices, which are recalled in this section. It is by using Mobile Robotics, First Edition. Luc Jaulin. \u00a9 ISTE Ltd 2019. Published by ISTE Ltd and John Wiley & Sons, Inc. this tool that we will perform our coordinate system transformations and position our objects in space. 1.1.1. Definition Let us recall that the j th column of the matrix of a linear application of Rn ! Rn represents the image of the jth vector ej of the standard basis (see Figure 1.1). Thus, the expression of a rotation matrix of angle ! in the plane R2 is given by: R = \" cos ! \" sin ! sin ! cos ! # . Concerning rotations in the space R3 (see Figure 1.2), it is important to specify the axis of rotation. We distinguish three main rotations: the rotation around the Ox axis, around the Oy axis and around the Oz axis. The associated matrices are respectively given by: Rx = $ % 1 0 0 0 cos !x \" sin !x 0 sin !x cos !x & ' , Ry = $ % cos !y 0 sin !y 0 1 0 \" sin !y 0 cos !y & ' , Rz = $ % cos " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002142_b978-0-12-816865-3.00010-x-Figure10.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002142_b978-0-12-816865-3.00010-x-Figure10.5-1.png", + "caption": "FIGURE 10.5", + "texts": [ + " Interfaces 2017;9: 30133e30142. https://doi.org/10.1021/acsami.7b08025. Copyright (2017) American Chemical Society. interference pattern, showing an exact replica of the irradiated light intensity pattern. Using two or multiple successive inscription steps of one-dimensional SRGs with grating vectors oriented in different directions [85e88], it is possible to obtain even more complex and periodic superficial texture on large scale. This multiexposure technique includes a rotation stage (see setup scheme in Fig. 10.5A) on which the azopolymer film can be mounted and exposed to the interference pattern for a fixed time. In each sequential exposure step, the sample is rotated to an a angle around the surface normal to achieve the required orientation of the new grating vector in respect to the previous one. The two-dimensional geometry of the resulting surface relief at the end of the multistep SRG inscriptions is given by the sum of the single one-dimensional SRGs inscribed with different grating vector in each step", + " (B) AFM image of the SRGs generated by rotating the azopolymers by 90 degrees in two subsequent irradiation steps and its Fourier transform. (C) AFM image of the SRGs generated by six-step subsequent irradiation (a \u00bc p 6) and its Fourier transform. rotation angles. In particular, symmetric ordered textures are obtained if constant rotation angles of p n are chosen for n sequential exposure steps [88], while inhomogeneous rotation amounts produce nonsymmetric two-dimensional textures on the surface. In Fig. 10.5B and C are reported, as example, AFM images collected from complex SRG obtained by means of the described technique for two-step and six-step, respectively. In this way, periodic surface reliefs, extended over areas of a few square centimeters, with amplitudes up to several hundreds of nanometers [89] and periods ranging from hundreds of nanometers to several micrometers can be easily obtained in very simple and cost-effective experimental conditions. All the methods described to obtain azopolymer textures with a degree of structural complexity higher than the sinusoidal SRGs are typically accomplished by serial illumination steps and are also limited in terms of versatility and complexity of the achievable surface textures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001428_012115-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001428_012115-Figure7-1.png", + "caption": "Figure 7. Fields of residual stress distribution over the section of the workpiece after transverse running with flat plates: a - intensity of residual stresses; b - radial residual stresses; \u0432 - tangential residual stresses; g - axial residual stresses", + "texts": [ + "05 mm, the working stresses are less than the yield strength \u03c3y (360 MPa) and therefore only elastic deformation is expected in this case. When the value of \u2206H is greater than 0.25 mm, the working stresses are greater than the tensile strength \u03c3t (600 MPa) and therefore, under such processing conditions, the material may be destroyed. Thus, the optimal value of the absolute compression is in the range \u2206H = 0.07\u20130.25 mm. Due to the Ansys Workbench program [18, 19], the results of calculating the residual stresses after transverse rolling at an absolute reduction \u2206H = 0.1 mm were obtained (Fig. 7). The stress intensity increases continuously from the center to the subsurface layers, and then decreases slightly on the cylinder surface (see Fig. 7.a). The radial residual stresses over the cylinder cross section (see Fig. 7.b) are tensile and monotonously increase from the surface of the cylinder to its center. The distribution of tangential and axial residual stresses is also alternating (see Fig. 7.c and d). For them, the maximum compressive stresses are observed at a certain depth from the periphery, and the maximum tensile stresses are observed in the cylinder central zone. \ud835\udf0e\ud835\udc64\ud835\udc5c\ud835\udc5f\ud835\udc58 \ud835\udc56 , \u041cP\u0430 ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012115 IOP Publishing doi:10.1088/1757-899X/632/1/012115 The experiments were performed on an experienced rolling machine. Cylindrical specimens with a diameter of 10 and a length of 100 mm from steel St45 were used as the investigated parts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000090_052079-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000090_052079-Figure9-1.png", + "caption": "Figure 9. Beams for use in grillages in slim floor technology: a) made from I-section, b) made from rectangular pipe.", + "texts": [ + " The cross-sections of beams can be Ishaped or rectangular, made from halves of I-sections and a flat bars, or in the form of a plate girder, or made by connecting an additional flat bar to a typical I-section or rectangular profile (figure 8). The implementation of the slim floor technology in grillages is possible, it only requires the use of sections in which the flat bar, which widens the bottom flange, is at either end shorter than the main section of the beam by half the difference between the widths of the flat bar and the section (figure 9). Such cross-sections enable the ends of one beam to be supported on the widened bottom flanges of the other beam (figure 10). Since the grillage is first loaded with its own weight, next with the weight of the slab and then with the live load, the connection may be reasonable at least during the assembly. Figure 11 shows a beam made from a rectangular tube with an added flat bar that widens the base and plates that increase the beam support surface area. The choice of the cross-section is justified by the creation of conditions for supporting the elements that fill the grillage boxes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003964_j.ijleo.2020.165806-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003964_j.ijleo.2020.165806-Figure5-1.png", + "caption": "Fig. 5. Data collection site of 222-type ODM system.", + "texts": [ + " Its structural parameters are shown in Table 1. The calibration test elements include 6 displacement sensors, sensor stand, and reference block to be measured (target reference block), etc., as shown in Fig. 4. The displacement sensors are SM30 series Czech ESSA grating displacement sensors, which have the advantages of large measurement range, high resolution and high accuracy, etc. [24]. Based on the 222-type ODM system, the calibration test was carried out, and the photo of the test site is shown in Fig. 5. The displacement sensors S1 and S2 were located in the + Z direction of the target reference block, S3 and S4 were located in the -Y direction, and S5 and S6 were located in the -X direction. The monitor 1 shows the data of S1, S2, and S3; and monitor 2 shows the data of S4, S4, and S6. Measurement configuration consisted of 18 positions and postures was selected, as shown in Table 2; the corresponding theoretical values of the 18 displacement sensors are shown in Table 3; displacement measurement was conducted, and the indicating values of the displacement sensors are shown in Table 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001999_14484846.2020.1714352-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001999_14484846.2020.1714352-Figure5-1.png", + "caption": "Figure 5. A conical helix trajectory followed by tool tip of the manipulator spindle.", + "texts": [ + " To this aim, a computer programme is developed in Matlab software, based on the algorithms outlined in Sections 7.3 and 7.4. To verify the mathematical results, dynamics of the manipulator is also simulated using SimMechanics toolbox of Matlab software. Architecture and mass properties of the manipulator are given in Appendix F. In the first example, the inverse dynamics of manipulator is analysed. Time trajectory of tool tip of the spindle is given, as follows h\u00f0t\u00de \u00bc 55\u00fe \u00f020= ffiffi 3 p \u00det cm \u03c6x\u00f0t\u00de \u00bc \u03c0 6 t 60 rad \u03c6y\u00f0t\u00de \u00bc 0 rad \u03c6z\u00f0t\u00de \u00bc t rad (80) where 0 t 8\u03c0 second. As shown in Figure 5, with these relations, tool tip of the spindle follows a conical helix trajectory of initial and final radii 10 and 2.09 cm, respectively. Moreover, feed of the spindle (or height of the helix) will be 290.208 cm. The SimMechanics model for the inverse dynamics of manipulator is presented in Figure 6. In this model, a motion subsystem is considered instead of the spherical joint of the PS leg which includes a bushing joint with six DOFs (Figure 7). The above time trajectory along with its time derivatives is applied on DOFs of the bushing joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002189_acit47987.2019.8991028-Figure14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002189_acit47987.2019.8991028-Figure14-1.png", + "caption": "FIGURE 14. Hydraulic principle of EHHT test system.", + "texts": [ + "13 shows the step signal response at amplitude 8MPa. Given that the load port pressure is about 0.15s delay at both control strategies, and considering the friction torque of EHHT, load pressure begins to increase when the EHHT\u2019s control angle increases and the driven torque is greater than friction torque. The simulation results show that the Fuzzy-PID controller can increase the EHHT\u2019s control performance. V. EXPERIMENTS We built a special laboratory bench for experiment of EHHT, the hydraulic system principle of EHHT is shown in Fig.14. The experiment bench is composed of following three parts. VOLUME 4, 2016 8613 (1) Main power of hydraulic circuit. This section provides a constant source of pressure for the electro-hydraulic servo plate-inclined plunger hydraulic transformer. It consists of a main motor, constant pressure variable pumps, hydraulic accumulators, safety valves, etc.. (2) Auxiliary oil source. This section provides controlled oil source for the laboratory bench. It mainly consists of an auxiliary motor, gear pumps, relief valves, hydraulic accumulators, hydraulic servo valve and filter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure29.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure29.6-1.png", + "caption": "Fig. 29.6 a Stress contours of cylindrical tool pin and b stress contours of taper cylindrical tool pin", + "texts": [], + "surrounding_texts": [ + "As these stresses play a very important role in the working performance of the component, it becomes quite necessary to examine the magnitude and nature of the residual stresses contained in the component. Stresses present can be tensile or compressive in nature but generally tensile nature of residual stresses is undesired as it can because quench cracking, stress-induced corrosion cracking, and fatigue failure. Compressive nature is desired one as it gives opposite effect of the tensile stresses on the component. The developed FE analysis model was validated by comparing the simulation of the turning process with the experimental data. Figures 29.5 and 29.6 show the simulation of turning process by using Abaqus 6.14 software. Due to advancement in technology and new methodology, now residual stresses due to turning operation can be visualized using computer simulation. Experiments have done by using Pulsetec\u00b5X-360n portable stress analyzer machine. Residual stress name is given to the stress which is left in the component after removal of the load. It remains almost in every rigid structure generated due to every operation performed on it such as metallurgical, thermal, and mechanical during manufacturing. Simulation of residual stress is an important process when H13 tool steel is used for CNC turning operation. Measurement of residual stresses due to CNC turning can now be visualized via computer simulations due to advancements in technology. 344 R. Butola et al. Unfortunately, turning processes have not yet been simulated to perfection in software due to a large number of deformations and variables involved. To overcome this problem, FE analysis is carried out in Abaqus/CAE 6.14. Experimental data is obtained from by using a Pulsetec\u00b5X-360n portable stress analyzer setup. Experimental and simulation results show the percentage error lying within the acceptable range for both the cylindrical and taper cylindrical tool of H13 tool steel (Figs. 29.7, 29.8, 29.9 and 29.10; Tables 29.3 and 29.4). The outcome of the residual stress on the periphery of a cylindrical pin and shoulder was observed to be compressive 166 and 280 MPa, respectively. The standard deviation was 16, 19 MPa, respectively. The outcome of the residual stress on the periphery of the taper cylindrical pin and shoulder was observed to be compressive 301 and 444 MPa, respectively. The standard deviation was 17 and 18 MPa, respectively. After the measurement of residual stress, an experimental data and simulation data for cylindrical tool pin were found to be compressive 223 MPa and compressive 240 MPa, respectively, so the percentage error was observed that 7.62%. For taper cylindrical tool pin, an experimental data and simulation data were found to be compressive 372 MPa and compressive 398 MPa, respectively, so the percentage error observed was 6.98%. 29 CNC Turning and Simulation of Residual Stress \u2026 345 346 R. Butola et al. 29 CNC Turning and Simulation of Residual Stress \u2026 347" + ] + }, + { + "image_filename": "designv11_80_0000384_scis-isis.2018.00219-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000384_scis-isis.2018.00219-Figure4-1.png", + "caption": "Fig. 4. An autonomous mobile robot (Crawler type)", + "texts": [ + " The search algorithm is shown in Fig. 3. Although it does not guarantee the converge to the optimum value by iterative search method, it seems to be sufficient for obtaining a reasonable route effectively by utilizing the path that passed on the outward way. III. EXPERIMENTAL CONDITION The specification of an autonomous mobile robot used in the simulation is set as follows. We assume that the robot moves on the rough terrain like a disaster site, the compact crawler type mobile robot is selected as shown in Fig. 4. The robot equipped with a laser range finder, and it is assumed that the external environment of 360 degrees can be acquired as threedimensional data. By updating the environmental maps, it quantifies the traveling risk evaluation points and possible to judge the dangerous area and plan the route autonomously by the installed computer. The functionality of a virtual mobile robot is tabulated in Table 1. The environmental map used for the experiment is virtually created as shown in Fig. 5, assuming a three-dimensional topography based on images taken with drone" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001185_10402004.2019.1664685-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001185_10402004.2019.1664685-Figure1-1.png", + "caption": "Figure 1. Schematic of a wet clutch: 1, Steel plate; 2, clutch drum; 3, output shaft; 4, oil cylinder; 5, piston; 6, return spring; 7, oil passage; 8, input shaft; 9, control passage; 10, transmission gear; 11, friction plate.", + "texts": [ + "RTICLE HISTORY Received 11 April 2019 Accepted 29 August 2019 KEYWORDS Wet clutch; fluid\u2013solid coupling; rub-impact; drag torque A wet clutch contains multiple pairs of friction pairs. According to the structure of the wet clutch (see Fig. 1), the friction pair and the spline are in mesh, the friction plates (FPs) are connected to the input shaft by the external splines with input torque from the engine, rotating around the Zaxis with the speed of w (rad/s), and the steel plates (SPs) are connected to the output shaft by internal splines. In this article, the dynamic behaviors of the friction pairs are studied when the wet clutch is in a braking condition, so the rotation speed of the SP is zero. Due to the clearance between the FP and SP and the assembly error gap between the friction pair and the spline, under the fluid action, the FPs and SPs will translate along the Z-axis with linear displacement z and oscillate around the X- and Y-axes with angular displacements a and b, respectively (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure72.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure72.3-1.png", + "caption": "Fig. 72.3 Meshed geometries of the diamond tip and the workpiece", + "texts": [ + " The quantity J2 is given by the following: J2 \u00bc 1 6 r1 r2\u00f0 \u00de2 \u00fe r2 r3\u00f0 \u00de2 \u00fe r3 r1\u00f0 \u00de2 h i \u00f072:2\u00de where ri are the principal stresses. The quantity j is given by the following, which is equal to the yield stress in the case, when rc \u00bc rt, that is, no pressure dependency: j \u00bc 2rcrt rc \u00fe rt \u00f072:3\u00de where rc and rt are the yield stress in tension and compression, respectively. The elastic materials properties for diamond and silicon are listed in Table 72.1 and Drucker\u2013Prager material constants for silicon are listed in Table 72.2. 72 Numerical Simulation and Experimental Validation \u2026 865 Figure 72.3 shows the meshed models of the diamond tip and the workpiece. In the present work, two-dimensional, four-node quadrilateral element CAX4R was employed to discretize the model geometries. It is four-node quadrilateral element and provides reduced integration with hourglass control. This element is chosen here because it is relatively inexpensive in terms of computation cost and time for problems involving non-linear constitutive behavior. For this type of element, the material calculations are only done at one point in each element by integrating all four nodes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure1.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure1.3-1.png", + "caption": "Fig. 1.3 AM production flow (Source Oak Ridge National Lab.)", + "texts": [ + " The industry accepted file format for AM is the STL, short for Stereolithography which was developed in the early 1980s. This format represents a computer aided drafting (CAD) model\u2019s geometry by faceted surfaces as shown in Fig. 1.2. This geometric model is virtually \u201csliced\u201d into layers and used to generate deposition paths for each layer of the component. Each layer is deposited sequentially on top of the previous layers to form the finished component. The production process flow is shown pictorially in Fig. 1.3. Support material is removed from locations with overhangs and finishing operations are performed tomeet the specifications on geometry, surface quality and/or resolution. Often these finishing operations involve sanding, vapor distillation smoothing, or machining. The benefits realized through AM are achieved through elimination of tools such as forging dies, reduction in production waste, creating functional structures, and in the production of components where traditional manufacturing operations are either prohibitively expensive, require significant tooling for limited production runs or not possible by traditional processes at all" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000316_978-981-13-6469-3_14-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000316_978-981-13-6469-3_14-Figure4-1.png", + "caption": "Fig. 4 Ball bearing geometry", + "texts": [ + " The analyzer essentially tries to pick up some picks, otherwise be buried in the high frequency noise. Envelope analysis steps are as shown in Fig. 1. The experimental test rig is as shown in Fig. 2. The test rig consists of a shaft with rotor disc and it was supported on two ball bearings. A DC motor coupled by a rigid coupling drives the shaft. For the vibration measurement and analysis, four channels FFT analyser is used (Fig. 3). The three piezoelectric accelerometers are used to collect the vibration signals in three directions at a bearing-housing. The ball bearing geometry is shown in Fig. 4. The self-aligning, double row deep groove ball bearing (SKF 1205 ETN) is used with the dimensions shown in Table 2. The fault in the Inner race, outer race and cage created using electro-discharge machining (EDM), the rectangular notch of 7.5 mm length, width 0.55 mm and depth 1.15 mm is created in outer race (Fig. 5). Similarly, rectangular notch of 7.5 mm length, width 0.5 mm and depth 1.15 mm, created in the inner race (Fig. 6), and rectangular notch of 6 mm length, width 0.7 mm and depth 1 mm, in cage (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003065_s00034-020-01490-y-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003065_s00034-020-01490-y-Figure6-1.png", + "caption": "Fig. 6 Responses of states x1 and x2 for case 2", + "texts": [ + " The fuzzy grade of h j is described as v ( h j ) = 1 \u03c3\u221a2\u03c0 exp ( \u2212(h j\u22121.5) 2 2\u03c3 2 ) , j = 1, 2, 3 with \u03c3 = 1 6 . Then, choose the parameter adaptive law (59), the intermediate virtual controller (30) and the actual controller (70). The parameters are designed as m1 = 4.6, m2 = 20.2, a1 = 3, a2 = 0.37, \u03b3 = 1.5, c1 = 0.3, c2 = 3 and k0 = 7. The initial conditions are defined as x1 (0) = 0.1, x2 (0) = 0.2, x\u03021 (0) = x\u03022 (0) = 0 and \u03c0\u0302 (0) = 0. The simulation results under Case 2 are shown in Figs. 6 and 9. The responses of states x1 and x2 are displayed in Fig. 6. The curves of control input u and fuzzy dead-zone \u03a0 (u) are given in Fig. 7. Figure 8 plots the curve of adaptive law \u03c0\u0302 . Figure 9 shows the error of y tracking yd . From the simulation results, it is obvious that even h\u0303 \u2208 [1, 2] is a fuzzy value, the effectiveness of the presented method can still be guaranteed. An observer-based adaptive NN control strategy has been presented in this paper for a category of nonlinear systemswith prescribed performance and fuzzy dead-zone input. The observer has been designed to estimate unmeasurable states, and the unknown nonlinear functions have been approximated by NNs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003966_032127-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003966_032127-Figure2-1.png", + "caption": "Figure 2. The sketch map of the robotic arm and the established joint coordinate system", + "texts": [ + " First, define the DH coordinate system on each joint to describe the current spatial pose of each joint, and use the homogeneous transformation matrix to obtain the posture relationship of the adjacent boom, and finally get the spatial pose transformation of each joint relative to the base coordinate system. Refer to the mechanical arm architecture designed in Figure 1 to simplify the components according to the connection method. The arm label from the origin to the end of the manned platform is 0,1,2 ... 7, the simplified model is shown in Figure 2, and the DH parameters are shown in Table 1. matrix of the coordinate of the i-th link to the i-1th link is shown in equation (1) Journal of Physics: Conference Series 1650 (2020) 032127 IOP Publishing doi:10.1088/1742-6596/1650/3/032127 \ud835\udc34\ud835\udc56 = [ cos(\ud835\udf03\ud835\udc56) \u2212 sin(\ud835\udf03\ud835\udc56) sin(\ud835\udf03\ud835\udc56) sin(\ud835\udefc\ud835\udc56\u22121) \ud835\udc4e\ud835\udc56\u22121 cos(\ud835\udf03\ud835\udc56) sin(\ud835\udf03\ud835\udc56) cos(\ud835\udf03\ud835\udc56) sin(\ud835\udefc\ud835\udc56\u22121) \u2212 cos(\ud835\udf03\ud835\udc56\u22121) sin(\ud835\udefc\ud835\udc56\u22121) \ud835\udc4e\ud835\udc56\u22121 sin(\ud835\udf03\ud835\udc56) 0 sin(\ud835\udefc\ud835\udc56\u22121) cos(\ud835\udefc\ud835\udc56\u22121) d\ud835\udc56 0 0 0 1 ] (1) The positive kinematics model of the seven-degree-of-freedom manipulator is obtained as shown in equation (2): (2) In this paper, the differential transformation method is used to calculate the Jacobian matrix of the manipulator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001886_summa48161.2019.8947534-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001886_summa48161.2019.8947534-Figure1-1.png", + "caption": "Fig. 1. Two-link manipulator kinematic scheme.", + "texts": [ + " Moreover, for manipulators with geometric constraints, the accuracy of the nonlinear model is of particular importance, since the stability of the steady-state motions of such manipulators is possible only in critical cases: the characteristic equation of the first approximation system with any control method necessarily has as many zero roots as there are geometric constraints imposed on the system. This greatly complicates the study of stability. II. FORMULATION OF THE PROBLEM Consider a two-linkmanipulator, rearranged in Fig. 1, controlled by two electric drives P1 and P2. At points O,B1,A1,B2,D,A2 cylindrical joints are installed. In ontrast to the traditional control of the rotation of the links of the manipulator due to the moments applied in the cylindrical joints O and D, a new method of realizing the force action using nonlinear geometric constraints is proposed here. Obviously, this design has significantly greater rigidity and, accordingly, significantly reduces the deformation of the manipulator links in comparison with the traditional case of the implementation of control actions directly in the rotary joints of the links" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001670_icems.2019.8922498-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001670_icems.2019.8922498-Figure10-1.png", + "caption": "Fig. 10. Equivalent Bearing Support Model", + "texts": [ + " The end cover of the motor and the rotating shaft are connected by bearings, but the internal structure of the bearing is quite complicated, including balls, inner and outer raceways, sealing covers and the like, that the calculation amount is greatly increased in the finite element analysis. In fact, the volume and mass of the bearing are very small in the overall structure of the motor. Therefore, in the modeling analysis, the bearing is often simplified to the equivalent joint stiffness and damping, instead of the solid model, as shown in Figure 10. In order to accurately analyze the influence of the bearing connection on the combined mode of the bearing and the end cap, the Y-90, 100, 132 combined modal test samples for making single shaft and end cover are shown in Fig. 11. The relevant parameters are shown in Table 9, and the modal parameters of the assembly are shown in Table 10 by free modal test. From the mode shape identified in Figure 12, the low-order modal frequencies are lower than others, mainly based on the rigid mode of the end cap and the shaft which means the bearing joint stiffness is very low" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000948_ever.2019.8813582-Figure20-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000948_ever.2019.8813582-Figure20-1.png", + "caption": "Fig. 20. Optimal machines with and without unequal rotor teeth.", + "texts": [ + " The optimal machine 2 is obtained by using the first selection criterion and the combination of rotor pole arcs is pole arc 1=14o and pole arc 2=21o. The second selection criterion is to find the combination of rotor pole arcs with the highest open-circuit DC winding induced voltage reduction ratio, whilst maintains at least 95% phase fundamental back-EMF of the optimal machine 1. The optimal machine 3 is based on the second selection criterion and the combination of rotor pole arcs is pole arc 1=18o and pole arc 2=17o. The optimal machines by unequal rotor teeth are shown in Fig. 20. As shown in Fig. 21, unequal rotor teeth is an effective technique for the reduction of the open-circuit DC winding induced voltage. The peak to peak values of the open-circuit DC winding induced voltages are 0.333 V, 0.015 V and 0.045V for the optimal machines 1, 2 and 3, respectively. As a consequence, the peak to peak induced voltage reduction ratios are 95.43% and 86.54% for the optimal machines 2 and 3, respectively. The peak to peak values of the cogging torques are 58.13mNm, 5.16mNm and 10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure20.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure20.2-1.png", + "caption": "Fig. 20.2 Finite element meshing of solution domain", + "texts": [ + " The fundamental governing equation of heat transfer or conservation of energy in 3D Cartesian coordinate is expressed as qC @T @t v @T @x \u00bc @ @x k @T @x \u00fe @ @y k @T @y \u00fe @ @z k @T @z \u00fe _q \u00f020:1\u00de where (x, y, z) is the coordinate system associated with the moving heat source. q; k and C refer to density, thermal conductivity and specific heat of the base material, respectively. T and t represents temperature and time variable, respectively. _q is the internal heat generation per unit time and per unit volume and \u2018v\u2019 is considered as the velocity of the plasma arc in welding direction (x-axis) or moving coordinate axis. The meshing arrangement of base plate is shown in Fig. 20.2. Very fine meshing is provided in the fusion zone area to capture thermal history precisely as it is exposed to a very concentric heat source where the temperature gradient becomes steep. While coarser meshing was generated in the area far from the heat flux to reduce computational cost, Continuum solid eight nodded brick elements (DC3D8 type) were selected in the thermal analysis for diffusive heat transfer. The double-ellipsoidal parameters like width (b) and depth (c) of the double-ellipsoidal heat source were directly taken from the experimental results", + " The length of front (cf) and rear (cr) quadrant were estimated as a function of velocity [20]. In dissimilar welding, the fusion zone becomes a mixture of both base materials which will produce various intermetallic compounds during the 236 A. K. Sahu and S. Bag solidification and will have an impact the joint efficiency. Therefore, to correlate the weld cooling rate with intermetallic phase formation precisely, the temperature dependent thermo-physical properties of fusion zone were considered as the average of both base materials which is named as mixed zone as shown in Fig. 20.2. Temperature distribution corresponding to S1 welding condition along with the moving heat source is shown in Fig. 20.3. A peak temperature of 2169 K is achieved at the weld centre line. The fusion zone (FZ) is defined by 1730 K which is the liquidus temperature of AISI 316L, i.e. maximum between the base materials [1]. And the heat affected zone (HAZ) was defined below 1547 K which is the solidus temperature of Inconel 718, i.e. minimum between the base materials [21]. The region between these two isotherm contours is considered as mushy zone where liquid and solid phase coexists together" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure83.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure83.5-1.png", + "caption": "Fig. 83.5 Load application and boundary condition for the adapter", + "texts": [ + " at z = 0 Ux = 0, Uy = 0, Uz = 0 and at z = H Ux, Uy, Uz 6\u00bc 0. \u2022 A uniform load is applied on the top nodes at z = H. \u2022 To have the interface, a metallic bulkhead at top and bottom of the adapter (lower and higher Diameter) steel is considered with properties as follows. Fig. 83.4 Elements in the model Design load, in this model, is applied for two types of analysis, i.e. static and buckling. Axial compressive load used for analysing the structure is 10 tons and the loading pattern is shown in Fig. 83.5. The model is fixed at the bottom and load is applied at the top. A load of 10 tons is applied on 280 nodes at the top. Then load per node comes out to be 357.14 N. In addition to static analysis, modal analysis was also conducted on the space adapter model (20 20) to study the vibrations and calculate frequency of the adapter. Modal analysis is used to calculate the natural frequency and mode shapes of a continuous structure. A continuous structure has infinite degrees of freedom, but finite element approximates with a finite number of degrees of freedom" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure83.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure83.2-1.png", + "caption": "Fig. 83.2 Development of the cone", + "texts": [ + "3 Configuration The conical adapter is designed with two circumferential stiffeners in addition to metallic stiffeners at top and bottom and 20 crossed axial stiffeners with varying helical angle. A total of 40 layers with 0.5 mm thickness are considered for analysis. The helical angles with fibre path variation cater for a height of about 300 mm with outer diameter to inner diameter ratio of three. The helical ribs of cross section 8 8 to 20 20 in multiples of four have worked out to cater the axial payload of about 50 kg with 10 g axial force. A typical 2D view of space adapter is shown in Fig. 83.2. 83 Modelling and Analysis of Composite Conical Space Adapter 987 The empirical relations for calculating the path, angle and radii are as given below (from Eqs. 83.1 to 83.6). r\u00a3 sin b \u00bc rh \u00f083:1\u00de h \u00bc \u00a3 sin b \u00f083:2\u00de \u00a3 \u00bc a a0 \u00f083:3\u00de Also, as per Clairut\u2019s equation R sin a0 \u00bc r sin a \u00f083:4\u00de a \u00bc sin 1 R sin a0 r \u00f083:5\u00de Substituting Eqs. 83.2 and 83.5 in Eq. 83.3, we get, h \u00bc 1 sin b sin 1 R sin a0 r a0 \u00f083:6\u00de The geometric parameters used for the design and modelling the conical space adapters serve as the inputs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003446_jae-200020-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003446_jae-200020-Figure15-1.png", + "caption": "Fig. 15. Distribution of heat flow, (a) the overall distribution of heat flow, (b) the heat flow of winding in detail.", + "texts": [ + "\u00a014c shows the temperature distribution of stator winding, with gradient distribution. As to rotor can and PMs, Fig.\u00a014d\u2013e demonstrates the temperature distribution, with consistent distribution, the intensified temperature areas occur in the rotor slot. As to rotor yoke, Fig.\u00a014f shows the temperature distribution, the higher temperature areas occur in connection between yoke and PMs and the lower temperature areas occur in end surface, due to the heat dissipation environment. The distribution of heat flow is shown in Fig.\u00a015. Figure\u00a015a demonstrates the overall heat flow distribution, the characteristics are as follows: (1)\u00a0As to overview, the direction of heat flow heat flow is from shaft to frame, due to heat dissipation environment; (2)\u00a0For the heat flow, the intensified heat flow area occurs in stator can, due to the considerably higher stator can loss. Figure\u00a015b shows the heat flow distribution of windings in detail, with the established coordinate axis that is consistent with that un cor rec ted pro of ver sio n Table\u00a09 Temperature value of each component under natural cooling Methods T (\u00b0C) Conventional network Improved network FE (3D) T fr 29.43 29.43 30.13 TSY 43.70 43.66 43.68 TST 112 115.5 115.63 TOC 188.8 188.5 188.57 Twi 114.2 113.4 113.42 T IC 171.82 172.48 172.57 TPM 170.61 170.6 170.58 TRY 169.15 168.9 168.5 TEW 114.2 112.2 established in part E, the characteristics are as follows: (1)\u00a0The intensified area of heat flow occurs between the stator yoke and winding; (2)\u00a0Heat flow between winding and teeth is less" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure23.11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure23.11-1.png", + "caption": "Fig. 23.11 Experimental setup", + "texts": [ + "10 confirms the failure region by showing where shear strain, von Mises stress, and shear stress become maximum. Results are taken at the end of 36 microseconds for different energy levels. A minimum and maximum von Mises stress of 341.6 and 382.6 MPa were found at the various voltage levels. A better groove filling is observed while using 4 kV that leads to resist higher torque load. Therefore, 4 kV is found optimum voltage for the given material system and geometrical parameters. The experiment is carried out using a setup shown in Fig. 23.11 which has a similar geometrical parameter to the FE model. The same geometrical and process parameters are used to conduct the experiment. Discharge energy generated using 280 G. T. Areda and S. D. Kore 4 kV is applied, and the radial displacement is measured to compare with the simulation. The radial displacement was measured using cross section of crimped sample and compared with simulation as depicted in Fig. 23.12. Measured gap between the tube and groove surface at the deformation zone was found in good agreement with the simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure10-1.png", + "caption": "Figure 10. The process of manufacture. The 3D printed model(a) is used as the pattern in the casting shaft(b), after the heating(c) the PLA model is disappeared and pour the ZL104 inside(d). The whole casting(e) is milled for the part(f).", + "texts": [ + "1088/1755-1315/332/4/042047 channel, which means the well sequence of solidification. Therefore, the best solution of gating system is cross-sharp side gating system with four risers. 3.2 Experiment and confirmation Mix the thermo stability plaster with water as proportion 100:46 and paint the plaster onto the 3D printed mould. Put it into the casting mold when the plaster is dried. Airing the whole mold for 3 hours and heating them in 400\u2103 for 15min then 600\u2103 for 30min. Make sure that the PLA mould is burned. Finally pour the ZL104 into the mold. The casting is displayed as Figure 10(e). The impeller is not only full filled with clear appearance, but without any defects. The post machining are cutting the gating system and milling the impellers. The final impeller is in Figure 10(f). (a) (b) ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 In summary, application of 3D printing mold in investment casting is experienced. The CAE analysis for advanced process of impeller is made and proved. In addition, the CAE analysis is also used for further improvements of solutions. The results are bellow. (1). The 3D printed mold can produce the qualified casting with excellent surface qualification and accurate size" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000293_012101-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000293_012101-Figure2-1.png", + "caption": "Figure 2. Maximum principal stress of (a) structural steel, (b) alloy steel, and (c) carbon steel.", + "texts": [ + " Maximum equivalent stress compared to metal yielding (Table 1) is a useful indicator to see the safety level of metals [11]. Maximum equivalent stress must be lower than the metal yielding [12]. From three materials tested, none of them met the safety criteria. However, to compare the three materials still can use other indicators through ANSYS simulations. IC2MAM 2018 IOP Conf. Series: Materials Science and Engineering 515 (2019) 012101 IOP Publishing doi:10.1088/1757-899X/515/1/012101 Maximum Principal Stress which specifically shows the most tensed part indicated by red color. Figure 2, below shows the maximum principal stress of structural steel, alloy steel, and carbon steel. The results of Maximum principal stress showed that alloy steel had 407.33 MPa, Structural steel had 411.37 MPa and Carbon steel had 408.74 MPa. The principal stress coefficient plays an important role in increasing strength [13]. This causes consequences on research\u2019s result, proven that structural steel has the highest strength because of the high value of principal stress. IC2MAM 2018 IOP Conf. Series: Materials Science and Engineering 515 (2019) 012101 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003983_icma49215.2020.9233584-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003983_icma49215.2020.9233584-Figure1-1.png", + "caption": "Fig. 1 The flexible body model of inner ring gear and planetary carrier", + "texts": [ + " The modeling process is described as follows: Firstly, a 3d rigid body model of planetary gearbox is constructed in the 3d modeling software SolidWorks, and gear parameters are shown in Table \u2160. Without affecting dynamic response of key components, the model is reasonably simplified to reduce unnecessary interference. For instance, planet gear and carrier are connected by alignment pin, bearings, gaskets and other parts, here it is simplified to a cylinder. Secondly, using ABAQUS software to create the flexible body models of inner ring gear and planet carrier, as shown in Fig. 1. Finally, the rigid body model is imported into ADAMS software, then constructed flexible body models of inner ring gear and planet carrier are imported to replace corresponding parts of the rigid model. The rigid-flexible coupling model is shown in Fig. 2. When gears are meshed with each other to transmit torque, strain will occur at teeth root of gear. Hence, the tooth root strain signals contain important information of gear meshing process, which can be used as the basis to judge whether the model parameter setting is reasonable or not" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000374_978-94-024-1620-6_11-Figure11.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000374_978-94-024-1620-6_11-Figure11.3-1.png", + "caption": "Fig. 11.3 (a) Schematic illustration of chemiresistive sensors composed of SWCNTs and metallosupramolecular polymers (MSPs). (b) Chemical structure of MSP. (c) Sensing device constituted by SWCNTs/MSPs bridging two gold electrodes. Monitoring of electric current (I(t)) under 0.1 V. (Reprinted with permission from [34]. Copyright 2017 American Chemical Society)", + "texts": [ + " Ishihara et al. [34], reported a study in which two different batches of tubes, metallic or semiconductor, were tested as a mixture and separated to study their effect into SWCNTs-electronic devices. They reported the conjugation of SWCNTs and metallosupramolecular polymer (MSP) to form a composite for detection of diethyl chlorophosphate (DECP), a highly toxic nerve agent simulant. In their study, they developed a chemiresistive gas sensor prepared by wrapping SWCNTs with a square-planar Cu 2+ \u2212based MSP (Fig. 11.3a). Addition of anthracene units to MSP, aligned by Cu 2+ configuration, ensured effective interaction between SWCNTs and MSP through \u03c0 \u2212 \u03c0 interactions (Fig. 11.3b), leading wrapping and debundling of carbon nanotubes. After exposure to electrophiles, conductivity increased due unwrapping of MSP (Fig. 11.3c) with a detection limit of 0.1 ppm DECP. As mentioned before, to develop and perform the sensor platform, they conveniently separated batches of as-synthesized SWCNTs, demonstrating that SSWCNTs exhibited the highest sensitivity to target analytes when wrapped with MSPs. On the other hand, the responses of M-SWCNT sensors were smaller but more precise in terms of discriminate another stimulus, like saturated water vapor. They concluded also that combination of both, semiconductor and metallic SWCNTs, could be a better option for sensing due its better response to saturated water vapor relative to pure S- and M- SWCNT/MSP sensors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002544_icit45562.2020.9067121-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002544_icit45562.2020.9067121-Figure4-1.png", + "caption": "Fig. 4. Diagram of external forces", + "texts": [ + " TABLE I PMSM VARIABLES Variable Description \u03c4e Electromagnetic torque \u03c6f Permanent iman flux Ld Inductance in the D axis Lq Inductance in the Q axis id Current in the D axis iq Current in the Q axis J Inertia B Coefficient \u03c1 Density p Number pair poles Ri Estator resistance \u03c9r Angular speed Increasing in available torque on the wheels of the electric vehicle, single-ratio transmissions are used at a maximum speed cost, depending on the speed of the vehicle [17]. Tr = rTe (9) where, r is the transmission relationship value, Tr is the torque commanded and the estimated torque Te . Among the various models in references [18], [19] and [20], a nonlinear model with four wheels presented in [18] is selected. Fig. 4 presents a simple longitudinal model that explains the action of external forces on the bodywork of the vehicle and the interface of the tire on the road. Equations (10), (11) and (12) describe this dynamic model that is determined by the Newton rules application: dV dt = Te rm \u2212 Faero + g(\u2212sin\u03b8 \u2212 ufcos\u03b8) (10) a) Aerodynamic force: This force is given by the following equation. Faero = 1 2m \u03c1CdA(V + Vm)2 (11) b) Rolling force: The rolling force uf acting on the wheel is linked to speed and its conversion factors: uf = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure15-1.png", + "caption": "Figure 15 Ninth order", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0001569_ecce.2019.8912739-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001569_ecce.2019.8912739-Figure9-1.png", + "caption": "Fig. 9 Prototype of the ASRM rotor", + "texts": [ + " The analytical equations of the total stress within the air gap bridge are summarized in Table II. To sum up, at this stage, there are two corrected analytic models that allow the determination of high speed mechanical stresses on the ASRM. III. NUMERICAL VERIFICATIONS In order to verify the accuracy of the proposed analytical methods, a FE model of the rotor is investigated. The rotor is built with the mechanical module in MotorCAD and with the OptiStruct solver from Altair. The main parameters of the prototype of the ASRM rotor (Fig.9) are listed in Table III. The stress values are highly sensitive to the iron bridge thicknesses and the rotational speed. Thus, in the following mechanical study, the MMS are calculated with respect to these two key parameters. Table IV gathers the mechanical characteristics of the rotor materials. The contact conditions between magnets and rotor laminations are defined at the interfaces shown in Fig.10. At these places, the rotor body and the magnets always stay in contact. No contact surfaces are configured in the other interfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000804_j.trpro.2019.07.074-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000804_j.trpro.2019.07.074-Figure2-1.png", + "caption": "Fig. 2 3D model of investigated rolling ball bearing", + "texts": [ + " \ud835\udf14\ud835\udf14 \u2219 (\ud835\udc61\ud835\udc61 \u2212 \ud835\udf0f\ud835\udf0f)\ud835\udc52\ud835\udc52\u2212\ud835\udc57\ud835\udc57\ud835\udc57\ud835\udc57\ud835\udc57\ud835\udc57 > (2) Where the symbol (*) matches complex conjugation and t - time shift of the window. The most commonly used window elements are the Hamming, Kaiser, Gauss or Hann rectangular windows [1, 4, 7, 8]. Peter Sulka et al. / Transportation Research Procedia 40 (2019) 511\u2013518 513 Sulka et al / Transportation Research Procedia 00 (2019) 000\u2013000 3 The aim of the analysis was to examine and analyze the oscillation of the investigated object, which was rolling ball bearing (see Fig. 2, 3) connected to the rotary shaft in the test center (see Fig. 4). The measurement itself was carried out in several steps to examine and compare the individual bearings with respective deformations in the orbits. The basis of the measurement was to obtain data on the dynamic behavior of individual bearings without and with deformations in orbits for the need to detect vibrations and to understand the behavior of the examined mechanical components during operation. Measurements were made at the start-up of the test device (with appropriate bearing on the rotary shaft) at 1000 rpm, staying at this level of mentioned rpm and runout of testing machinery in a total cycle of approximately 50 s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000055_012044-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000055_012044-Figure4-1.png", + "caption": "Figure 4. Pressure contours without control (\u03c9=7000rpm). \u00a0", + "texts": [ + " Series: Materials Science and Engineering 473 (2019) 012044 IOP Publishing doi:10.1088/1757-899X/473/1/012044 Air flow around the gear teeth has been investigated numerically. The pressure distributions on the tooth surface have been indicated to understand the mechanism of the windage power loss. The air is sucked into the gap between two teeth from both sides of the gear. Then it is thrown out of the gap along the radial direction. The pressure contours at gear and symmetry surfaces without control are shown in figure 4. It is shown that adjacent tooth surfaces have different pressures. The pressure difference induces the rotating resistance torque of high speed gear. To keep constant rotating speed, the corresponding power has been taken to overcome the resistance torque. A smooth shroud is applied for enclosing the high speed gear and restricting the pumping action by the gear teeth. The volumetric flow rate around the gear is reduced by enclosing the gear. The pressure difference is much lower with the smooth shroud, as shown in figure 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure2-1.png", + "caption": "Figure 2 Meshing result", + "texts": [ + " In order to ensure the quality of meshing and calculation accuracy, some simplifications are made when importing the 3D model. According to the actual size of the model, the selected grid cell size is 5mm. The statics analysis module is established on the ANSYS Workbench software platform, and then the above model is exported from Solidworks to the .x_t format and imported into ANSYS Workbench. The unit type used for meshing is solid 186, and the unit is selected (mm, t, N, s, mv, mA). The final result is shown in the figure below (Figure 2). As a result of the division, the number of nodes is 57124, and the number of units is 31436. ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 Setting the material parameters in the material property module an deselect the HT250 structural steel material that has been set. The main physical property parameters of this material are shown in the following table (Table 1): Disc brakes use the inner and outer caliper bodies as the main application objects during the braking process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002580_01691864.2020.1751707-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002580_01691864.2020.1751707-Figure1-1.png", + "caption": "Figure 1. An unexpected collision occurred between an n-DOF robotic manipulator and an unknown obstacle.", + "texts": [ + " An adaptive collision detection and isolation scheme was presented for robotic manipulator based on the generalized momenta, and a suitable overparameterization was considered to uncertain robot dynamics in [28]. To appropriately respond to a collision, several reaction strategies should be applied to robotic manipulator in the post-contact phase, such as stop immediately [29], move away from the obstacle [30] or continue the original task [31]. We consider an unexpected collision occurred at a point Pc on Link i of an n-DOF manipulator with an obstacle when the manipulator is performing a reaching task, as shown in Figure 1.The dynamic equation of the n-DOF manipulator is denoted as follows: H(q)q\u0308 + C(q, q\u0307)q\u0307 + g(q) + Fvq\u0307 + Fs sign(q\u0307) = u + JTc,i(q)f c, (1) where H(q) \u2208 R n\u00d7n, C(q, q\u0307)q\u0307 \u2208 R n and g(q) \u2208 R n are the inertia matrix, Coriolis and centrifugal vector and gravity force, respectively, Fvq\u0307 and Fs sign(q\u0307) are the viscous and static frictions, respectively, u \u2208 R n is the commanded torque, Jc,i(q) \u2208 R 6\u00d7n and f c \u2208 R 6\u00d71 are the Jacobian matrix and external contact force at the contact point Pc, respectively. The manipulator is not equipped with any force sensors, thus Jc,i(q) and f c can not bemeasured directly. It is well known that the left-hand side of manipulator dynamics (1) can be expressed as a linear function of link parameters as follows: H(q)q\u0308 + C(q, q\u0307)q\u0307 + g(q) + Fvq\u0307 + Fs sign(q\u0307) = Y(q, q\u0307, q\u0308) , (2) where Y(q, q\u0307, q\u0308) is a manipulator configuration-based regressor, and is a constant vector containing manipulator link parameters [32]. As shown in Figure 1, the unexpected collision might not only destroy the robotic manipulator but also be harmful for the unknownobstacle, especially in the applications of human-robot cooperation in which humans can be treated as the obstacles. Therefore, in order to prevent further damages of both manipulator and obstacle after the unexpected collision, we should know the contact information, e.g. the exact contact point or at least a specified contact link, because the manipulator could purposefully be responded to the unexpected collision according to a feasible reaction strategy that is designed using the contact information. In [19, 20], we proposed a set of instantaneous powerbased indexes to detect and isolate a link of a manipulator colliding with an unknown obstacle. As shown in Figure 1, our idea is to introduce a virtual contact point Pvi on each link of the manipulator and the corresponding virtual contact force is f vi \u2208 R 6\u00d71. These virtual contacts act as a series of virtual sensors equipped on the manipulator link individually; and we observe a set of instantaneous power-based indexes for collision detection and isolation, because the contact link can be distinguished according to the differences of impulsive changes of the power indexes, which are caused by the changes of velocity and force at each virtual contact point Pvi during the transition from non-contact to contact situation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000387_978-981-10-4938-5_13-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000387_978-981-10-4938-5_13-Figure5-1.png", + "caption": "Fig. 5 The principle of full melting of a workpiece", + "texts": [ + " Here the particles shown merge during the 218 sintering process, and neck growth occurs between the former separate metal 219 particles. Finally, shrinkage occurs, by which, during the neck growth, the center 220 parts of the particles move closer together and the compaction of the particles into a 221 solid makes the total volume shrink as compared to the original state (Kalpakjian and 222 Schmid 2010). 223 Melting 224 As opposed to sintering, the consolidation of the powder particles by melting is done 225 in a way that ensures that the targeted particles are fully molten, as seen in Fig. 5. 226 This approach is more difficult to control since minute changes in heat flux from the 227 heat source affect the size of the melt pool, affecting overhanging structures and in 228 general the down-face of the manufactured component. Whereas sintering within 229 AM is not exclusive to metals, characteristics of metal alloys such as conductivity, 230 and a distinct melting point, make them de facto exclusively used as materials for full 231 melting in powder consolidation systems. The rapid melting and solidification 232 enabled by the process allow for parts with tailored properties distinct from those 233 obtained from traditionally processed parts (Gibson et al", + " This is achieved by 234 the exploitation of localized material properties, obtained from refined microstruc235 tures, formation of non-equilibrium phases, and supersaturated solutions. Second236 phase particles such as inclusions and carbides, or extremely fine, refined micro237 structures, are side effects of the process. 238 Due to the similarity of powder-based sintering and melting, whether the 239 technology is referred to as sintering or melting, often both of the states are present. 240 Given that the sintering temperature is slightly below the temperature at which a melt 241 pool is formed, as shown in Fig. 5, sintering will occur in the boundary of a geometry 242 made from a melt-driven process. Similar to this, localized melting can occur during 243 a sintering process, where impurities and process fluctuation may temporarily lead to 244 the formation of melt. Melting is often preferred over sintering given that, in order 245 for the process to be as economically competitive as possible, the fusion of the 246 powder should happen as quickly as possible in order to increase the build rate. 247 This is achieved by a fast-moving energy source, with a high-energy flux, and fusion 248 by melting" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001930_vppc46532.2019.8952232-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001930_vppc46532.2019.8952232-Figure3-1.png", + "caption": "Fig. 3. Temperature distribution in SRM", + "texts": [ + " When keeping the rotate speed at 500 r/min, the machine iron loss calculated by the iron loss model at different torque load is shown in Fig. 2(b). IV. TEMPERATURE FIELD ANALYSIS AND TEST RESULTS On the basis of SRM iron consumption calculation model, A 2D finite element thermal model of the machine under NNNN-SSSS magnetic pole distribution is established through FLUX software and the temperature field model of machine is further analyzed. The temperature distribution calculated by 2D FLUX in SRM is shown in Fig. 3. It is shown that the highest temperature of the machine is located in the winding of the stator. Temperature of stator tooth is lower than that of the winding, but higher than that of the rotor. Two reasons can explain this situation. In one hand, iron loss of stator tooth is relatively larger. On the other hand, heat generated from winding can also be transmitted directly to the stator tooth. The temperature of rotor is the lowest, the reasons are as follows. One reason is that the iron loss of rotor tooth and rotor yoke is small" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001962_j.mechatronics.2020.102323-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001962_j.mechatronics.2020.102323-Figure2-1.png", + "caption": "Fig. 2. 1D Point mass", + "texts": [ + " Hence, \u2022 If the state x i has fixed value at boundary, BC (20a) is applied \u2022 If state x i is set to be free at boundary, BC (20b) is applied. The justification for the use of these boundary conditions can be btained following the Lagrangian approach developed in [21] More pecifically, the free boundary conditions can be obtained by applying ransversality condition on the Lagrangian. We now illustrate the use of these boundary conditions on what is erhaps the simple example for which they matter: a point mass moving n vertical direction under gravity. For a point mass with only gravity nd no control \u2014free falling, as shown in Fig. 2 , the state space can be efined as \ud835\udc65 = [ \ud835\udc66, ?\u0307? ] where y is the height of the point mass. We seek for he motion of the point mass which starts from ground with duration T . his corresponds to the boundary conditions (0) = 0 . e explore the effect of setting the remaining boundary conditions ?\u0307? (0) , ( T ) and ?\u0307? ( \ud835\udc47 ) : Case 1: ?\u0307? (0) , y ( T ) and ?\u0307? ( \ud835\udc47 ) are fixed. This corresponds to fixing initial ertical velocity, final height and velocity. Because the in-air motion onserves energy, the initial boundary conditions can be thought of as xing the value of said energy" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002751_s00162-020-00532-0-Figure14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002751_s00162-020-00532-0-Figure14-1.png", + "caption": "Fig. 14 The pressure distribution on the plane of xoy under the condition ofMa = 3 and \u03b1 = 6\u25e6", + "texts": [ + "15D boattail, a jump change of aerodynamic coefficients occurs, which is completely different from the monotonic change of the aerodynamic coefficients at angle of attack \u03b1 = 0\u25e6. Therefore, it is necessary to further explore the mechanism. F ig .1 3 T he pr es su re co nt ou r ar ou nd bo at ta il pa rt un de r th re e di ff er en tb oa tta il an gl es Firstly, the projectile with the boattail structure of lbt = 0.25D, \u03b8bt = 9\u25e6 is selected as typical example to compare with the original projectile with non-boattail. Figure 14 shows the pressure distribution on the plane of xoy under the condition ofMa = 3 and \u03b1 = 6\u25e6. Figure 15 shows the instantaneous Mach number contours of two kinds of projectiles. As seen from Fig. 14, the overall flow field structures of the two projectiles are similar, but the pressure distribution in the circular area of the warhead site and the tail position of the projectile is slightly different. The pressure difference of wake field is easy to understand, because it changes from a primary expansion wave to two expansion waves (as shown in Fig. 16). The mechanism for the evolution of the high-pressure area on the warhead is relatively complex. In Fig. 15, the velocity near the windward side increases due to the change in the boattail structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002992_icusai47366.2019.9124797-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002992_icusai47366.2019.9124797-Figure1-1.png", + "caption": "Fig. 1. Helicopter attitude simulation system", + "texts": [ + " In Section 3, a PID based controller with prescribed performance function is discussed. Comparative simulations for the helicopter attitude control by the proposed control Foundation items: National Natural Science Foundation of China (51809138). 978-1-7281-5859-4/19/$31.00 \u00a92019 IEEE 61 Authorized licensed use limited to: Carleton University. Downloaded on July 25,2020 at 23:22:27 UTC from IEEE Xplore. Restrictions apply. scheme are introduced in Section 4. Conclusions and future work are given in Section 5. II. PLANT MODELING Fig. 1 gives a helicopter attitude simulation system, which is a simplification of the common helicopter system. It removes the position motion of the helicopter and only contains the attitude movement. The multi-input- multioutput, non-linear and strong coupling characteristics in the attitude motion are fully preserved. The platform can be used for simulation and experiment of helicopter attitude control conveniently. First of all, two reference frames are defined to describe the attitude of the helicopter system", + " The other one is the body frame ObXbYbZb, where Ob is the intersection of the short rod and the long rod, the Xb-axis coincides with the long rod and points to the motor side, the Yb-axis coincides with the short rod and points to the right motor, and the Zb-axis perpendicular to the plane of XbObYb and points downwards. Before mathematical modeling, we make the following assumptions. 1) This helicopter system is rigid and symmetrical such that the moments of inertia Jxy, Jyz, Jxz are zero. 2) In the Fig.1, the gravity action lines of the short rod in pass through the roll axis respectively, so the gravity of the rods does not affect the roll attitude motion. 3) The mass center of the long rod is located at pitch axis, and the gravity of the long rod does not affect the pitch attitude motion. 4) The lift generated by the motor is always perpendicular to the plane of XbObYb, and the motor cannot be reversed to produce negative lift. 5) The Jz is large enough that the yaw interference torque caused by the different speed of two motors can be ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000635_978-3-030-20377-1_1-Figure1.20-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000635_978-3-030-20377-1_1-Figure1.20-1.png", + "caption": "Fig. 1.20 Receding contact", + "texts": [ + " The size and position of the permanent stick zone depend only on the contacting profile within the stick zone, and the overall size of the contact does not appear in the solution. Far fewer general results are available for the remaining two types of contact, and we provide, here, some example results. Before any external forces are applied, receding contacts conform, in the same way that complete contacts do. In some cases the initial extent of the contact is defined by the limits of conformality, such as the thin rectangular strip resting on the elastically similar half-plane shown in Fig. 1.20. Indeed, if the contact pressure is applied over a significant fraction of the upper surface, the contact may well extend along the whole of the lower length of the rectangle and not recede, but be complete. If, on the other hand, pressure on the upper surface stops well short of the ends they will tend to lift off, and so the contact will recede. Receding contacts of this kind have the property that the application of an infinitesimal load causes the contact to \u2018snap\u2019 to a reduced length, and an increase in the applied pressure does not cause a change in the contact length, so that the problem becomes linear, in the sense that the magnitude of the stresses and displacements (including the lift-off angle) are simply proportional to the applied load", + " With this modification, when the contact lifts (at a very small load) the number of node pairs where separation occurs is much smaller than if the gap was not present. Finite element programmes do not cope well with the sudden separation of a large length of the interface and this helps considerably. The tendency for a strip to lift off is very strongly controlled by the presence of friction along the interface. If we have a layer of thickness c resting on an elastically similar half-plane, we can think of the pair, together as a half-plane of additional \u2018depth\u2019 h, and, if a line load is applied, Fig. 1.20, use the Flamant solution to discover what happens. We can convert the solution derived in section\u201cHalf-Plane Problems\u201d into Cartesian coordinates, and this then gives p(x) \u2261 \u03c3yy (x, h) = \u22122Ph3 \u03c0 ( x2 + h2 )2 , (1.80) q(x) \u2261 \u03c3xy (x, h) = \u22122Pxh2 \u03c0 ( x2 + h2 )2 , (1.81) so that, prima facie, it would seem that intimate contact will be maintained between layer and substrate. But notice that, if we take the traction ratio \u03c3xy \u03c3yy = x h (1.82) keep h constant and increase x we can see that the traction ratio becomes infinitely large, and it would therefore require an infinitely large coefficient of friction to inhibit all slip", + " These are very effective at permitting discrete regions of slip and separation whilst continuity of material is maintained elsewhere. The solution is known for the state of stress induced by an edge dislocation in a half-plane, in closed form. For the purposes of the present problem, it is best if the origin lies on the surface of the true half-plane, on top of which lies a layer of thickness c, so that the \u2018augmented half-plane\u2019 (the actual half-plane and layer together) occupies the region y < c, Fig. 1.20. The tractions arising on the putative interface line, y = 0, generated by a dislocation, also on the line, installed at point \u03be and having Burgers vector component (bx , by) are given by { \u03c3xy (x, 0) \u03c3yy (x, 0) } = E 4\u03c0 ( 1 \u2212 \u03bd2 ) [ 1 x\u2212\u03be + Gxxy(x, \u03be) Gyxy (x, \u03be) Gxyy (x, \u03be) 1 x\u2212\u03be + Gyyy (x, \u03be) ] { bx (\u03be, 0) by(\u03be, 0) } (1.83) in plane strain, where the functions Gi jk account for the presence of the free surface and are given explicitly in (Chaise et al. 2014). Note that there are Cauchy singular terms\u2014the glide dislocation generates a singular behaviour in the shear traction and the climb dislocation develops a singular behaviour in the direct traction, but these singularities will integrate out when we have distributions of dislocations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000841_iceee2019.2019.00065-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000841_iceee2019.2019.00065-Figure6-1.png", + "caption": "Figure 6. Axes and direction of movement of quadrotor", + "texts": [ + " Control setpoint values are desired latitude and longitude and feedback value is current latitude and longitude information. Equation 25 and 26 illustrates this distribution where c is heading, R and P is uncorrected roll and pitch control outputs. Control coefficients are selected as 2.5 for P and 6.5 for D. Cascade control output signals define motion required for each axis. Control signals must be distributed to related motors. Direction of motor rotations and propellers are important in signal distribution. Figure 6 shows the axes, direction of motion and motor numbering. V. CONCLUSION This paper presents flight controller design based on STM32F407 microcontroller. On this base, Flight controller which can control the quadrotor both manually and autonomously, is realized by integrating microprocessor and peripherals. Flight tests show that flight controller can follow flight instructions rapidly and attitude of quadrotor is stable. The follow-up work is to integrate fuzzy controller for better performance in the case of different flight modes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000297_kem.799.263-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000297_kem.799.263-Figure7-1.png", + "caption": "Fig. 7 KLP-P-002 redesign", + "texts": [ + "90) KLP-P-009 Compression 1.55 / (-1.54) KLP-P-008 Shear stress 2.11 Min: -7.90 MPa Max: 1.94 MPa Fig. 6 Traction stress results on KLP-P-009 The proposed improvements follow the order of problem areas shown in Table 2. As such, the first area to improve is mobility between the crank webs. Here, the mechanical analysis showed that the LCD is not able to clad the entire surface and identified a first collision when the laser cladding nozzle swings between the nozzle support and the sidewall component. As can be seen in fig. 7, on the left of the image, the sidewall component of the guidance structure acts as a barrier that could be useful to prevent collision with the crank-web, but it also drastically reduces the swing mobility. To get an idea of the limitations, the original LCD could swing \u00b17 degrees. By modifying the component, as shown on the right-hand side of fig. 7, the swing motion increases up to \u00b115.5 degrees. Also, two additional sliding spindles (KLP-P-004) on each side should be included in the new LCD, giving a total of four spindles. Furthermore, the rack-pinion mechanism was moved from the bottom to the centre, between the two sliding spindles on each side (see central and right pictures in Fig. 7). The latter modifications had a direct impact on improving problem five, listed in Table 2. The second area of improvement was the nozzle-support structure. This area is particularly critical due to the dynamic and static efforts involved here. To begin with, the lifting movement needed to be improved in order to resolve the misalignment and blocking issues. To do this, the shape of the sliding component was submitted to a tolerance adjustment redesign and the material type was changed, reducing the blocking risk and friction, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure6-1.png", + "caption": "Figure 6. The solidification(a) of model with bottom gating system is unsymmetrical, and the combined defect (b) is appears on the blade. Probabilistic defect (c) is forecasted.", + "texts": [ + " Thus the surface of fluid is unsmooth which means the appearance of turbulence. The final results of filling can be explained by Figure 5(b) and Figure 5(c). The combined defect parameter analyses the possibility of defect while the probabilistic defect parameter proof that the shrinkage defects and gas porosity will occur between the gating system and the impeller like the zone printed in yellow. In bottom gating system, the metal fluid is poured through the bottom of cavity. The solidification time is shown as Figure 6, which means that the solidification of impeller is in an orderless situation. Many defects located in the blades according to the combined defect parameter. Moreover, like the top gating system , the shrinkage defects and gas porosity stay in the similar place. The reason of defect is suspected to be the high speed of filling, which leading to a phenomena that not all the gas escape through the riser on time. ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003311_speedam48782.2020.9161923-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003311_speedam48782.2020.9161923-Figure6-1.png", + "caption": "Fig 6. Calculated magnetic field strength distribution of PMaSynRM after short-circuit from motor no-load condition at t = 13.7 ms (JMAG Designer) (compare Fig. 5)", + "texts": [ + " The PMaSynRM is operated with nominal voltage UN = 400 V at nominal frequency fN = 50 Hz without any load. Only a magnetizing d-current and a small negative q-current to compensate the PM torque are needed (id0 = 8.2 A, iq0 = -1.4 A). By using the synchronous inductances from load operation and the PM flux linkage from generator no-load operation, the maximum values of short-circuit current and torque are again underestimated (Table III). These deviations can be decreased by a lower q-axis inductance, too. In order to meet the smaller first peak value of the short-circuit current (Fig. 6), a parameter study yields that the PM flux linkage has to be slightly increased in the analytical calculations. Fig. 6 shows the magnetic field strength at t = 13.7 ms, when the maximum short-circuit current occurs. The 100 Authorized licensed use limited to: Carleton University. Downloaded on October 05,2020 at 07:40:32 UTC from IEEE Xplore. Restrictions apply. magnetic field is again aligned near the q-axis, so that the maximum short-circuit current is also mainly a q-axis current. The magnetic field is therefore in opposite direction to the magnetization direction of the ferrite magnets. The calculated field strength values in the outer two magnet slots are above the knee point of the used ferrite magnets (250 kA/m), which leads to irreversible demagnetization there", + " by 30% at motor no-load of the SynRM) and the short-circuit torque (e.g. by 35% at motor no-load of the SynRM) is improved (Table II and III). Using typical motor parameters of SynRM and PMaSynRM, the transient short-circuit phase current has no zero crossings for several electrical periods, which is verified by measurements (Fig. 7). The transient short-circuit current is high enough and excites a magnetic field in opposing direction to the ferrite magnets of the PMaSynRM, so that they partially demagnetize irreversible (Fig. 6), which is confirmed by measurements (65% lower steady-state shortcircuit current). After that short-circuit, the operation of the PMaSynRM is still possible and investigated at the test bench. The output torque is reduced (e.g. by around 10% at an rms phase current of 20 A) as well as the maximum steady-state output power (e.g. by around 22% at 3000 rpm). [1] B. Janjic, S. Laue and J. Schaab, \"State estimation and process optimisation for multi-pump systems with synchronous reluctance motors,\" 2015 IEEE 10th International Symposium on Diagnostics for Electrical Machines, Power Electronics and Drives (SDEMPED), Guarda, 2015, pp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003440_aeat-01-2020-0015-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003440_aeat-01-2020-0015-Figure1-1.png", + "caption": "Figure 1 Quadrotor reference frames and Euler angles with respect to inertial frame", + "texts": [ + " While first pair (1 \u2013 3) rotates counter-clockwise, second pair (2 \u2013 4) rotates clockwise. Vertical motion is generated by increasing or decreasing the rotation speeds of all rotors, simultaneously. Pitch or roll motions are generated by difference between first and third rotors or second and fourth rotors, respectively. Yaw motion is realized by difference between the rotor pairs. As quadrotor is underactuated, X and Y states are controlled by pitch and roll motions, respectively. A quadrotor is given in Figure 1 with its associated reference frames, rotors and Euler angles. Two different frame systems are required to derive the kinematic equations of the quadrotor. First frame is body \u2013 fixed frame, denoted as B and its elements are {xb,yb,zb} [ R3. Second frame is earth \u2013 fixed frame, denoted by I and its elements are {xi,yi,zi} [ R3. Each rotor produces an upward force vector which is called as Fi where i = {1,2,3,4}. Linear and angular velocity vectors of the quadrotor that are defined on any frame can be converted to another frame via rotation matrices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002638_pesgre45664.2020.9070716-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002638_pesgre45664.2020.9070716-Figure9-1.png", + "caption": "Fig. 9. Flux density for 40 W (16 Pole) SPIM under no load condition", + "texts": [ + " The core is highly saturated under the no load condition as, significant magnetization current is drawn. Under locked rotor condition, the rotor resistance is comparable to the stator resistance. Hence, the current drawn from supply majorly heats up the stator (both the windings) and rotor resistance. Fig. 8 shows the flux density plot for 40 W SPIM motor under locked rotor condition. The stator and rotor core are unsaturated as seen in the plot confirming with 2.5386 W core losses. The rest of the losses are the copper losses (57.10 -2.5386 W= 54.5614 W). Fig. 9 shows the flux density plot of the 40 W SPIM motor under no load condition. The stator as well as rotor core are highly saturated confirming higher losses. The performance of the fan at desired speed (for a particular blade) is calculated from FEA. Moreover, the performance of the motor at various blades, is examined using the transient mechanical simulation. Fig. 10 shows the torque vs speed characteristics with four different blades. There is significant torque ripple at steady state. With the higher angle blade, the steady state speed is lesser, and torque is higher", + " 12 depicts the winding currents and supply current for various blades as mentioned in Table I. The flux density distribution of a 40 W ceiling fan motor at rated load is shown in Fig. 13. The rated power consumption for a 40 W (16 Pole), 50 W (14 Pole) and 50 W (12 Pole) SPIM motors are shown in Figs. 14, 15 and 16, respectively. Fig. 13. Flux density distribution in steel part of a 40 W SPIM at rated load The stator teeth are much less saturated as compared to the no load condition (as shown in Fig. 9). The core loss at rated load is 4.57 W. TABLE I. SIMULATED PERFORMANCE OF 40 W SPIM MOTOR WITH DIFFERENT BLADES (MULTI SLICE 2D FEA) Speed (rpm) Torque (Nm) Pout (W) Pin (W) Imain (mA) Iaux (mA) Isupply (mA) Stator Copper Loss (W) Rotor Loss (W) Core Loss Stator (W) Core Loss Rotor (W) Stray Load loss (W) Efficiency (%) Power Factor 317.41 0.52 17.45 43.13 152.63 223.04 192.53 11.44 3.12 4.73 0.42 5.97 40.45 0.974 314.67 0.54 17.90 43.78 153.97 221.97 195.09 11.48 3.22 4.68 0.42 6.08 40.88 0.976 309" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003112_j.mechmachtheory.2020.103903-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003112_j.mechmachtheory.2020.103903-Figure10-1.png", + "caption": "Fig. 10. Screenshot, from digital simulation in video file Xactuator.mp4 in supplementary materials, of extensions to coupled-Cartesian-manipulator family.", + "texts": [ + " Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx 5 Table 1 ( continued ) Symbol Description \u03b8 = atan 2( y, x ) Four quadrant arctan ( y/x ) function with range 0 \u2264 \u03b8 \u2264 2 \u03c0 P, R, U , C , S, P a Passive joints: prismatic, revolute, universal, cylindrical, spherical, parallelogram P , R Active joints: prismatic, revolute P 2 Redundant prismatic P joints ( \u00b7 ) Joint topology: parentheses ( \u00b7 ) enclose the joints \u00b7 of serial kinematic linkage (limb) ( \u00b7 ) ( \u00b7 ) Joint topology: two limbs ( \u00b7 ) connected together in parallel n ( \u00b7 ) Joint topology: n serial limbs ( \u00b7 ) connected in parallel 6 P. Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx P. Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx 7 Table 2 ( continued ) Manipulator References Joint topologies Features Coupled-Cartesianmanipulators with prismatic P joint along common-link L T between coupler-links L Bn , L Dn to enable changes in distance | z T Bn \u2212 z T Dn | Fig. 10 ( P PP U ) ( PP UP ) \u2217 Extension to family of coupled-Cartesian-manipulators to investigate in further research \u2217 By contrast, manipulators reported here have fixed distance between coupler-links L Bn , L Dn and prismatic P joints between positioner-links L An , L Cn in \u02c6 Z W W direction to accommodate changes in distance | z W An \u2212 z W Cn | as the common-link L T orientation changes 6-DOF coupled-Cartesianmanipulators with active revolute R joints for control of angles \u03b8A n Z , \u03b8C n Z or \u03b8 T Z Fig. 10 ( R P PP U ) ( P P P U ) \u2217 Extension to family of coupled-Cartesian-manipulators to investigate in further research \u2217 Enables rotation control of angle \u03b8T Z around common-link L T longitudinal \u02c6 Z T axis and D) In Sm illustrate manipulators with intersecting-revolute-axes. Each figure illustrates a stand-alone manipulator, with its own unique features that are summarized in Table 2 . The progression of Figs. 1\u20136 illustrates the hierarchical composition of the 5-DOF 3 T 2 R fully parallel mechanism of Fig", + " In this case the distances between positioner-links L An , L Cn in the \u02c6 Z W W direction may be fixed, i.e., they may have fixed position coordinates, z W An , z W Cn . A coupled-Cartesian-manipulator may be extended from 4-DOF 2 T 2 R to 5-DOF 3 T 2 R motiontype with an active prismatic P joint controlling either position coordinate z Bn T or z Dn T . Furthermore, the family may be extended to 5-DOF 2 T 3 R or 6-DOF 3 T 3 R motion-types with an active revolute R joint controlling either angle \u03b8 A n Z , \u03b8 C n Z or \u03b8 T Z as simulated in video file Xactuator.mp4 ( Fig. 10 ). The manipulator family may also be extended with additional limbs connected in parallel to the ones reported here. Kinematic analysis. The mathematical framework for the manipulator family can be extended to include the direct kine- matics equations for both non-intersecting and intersecting-revolute-axes cases. Closed-form Eq. (38) expresses the parasitictwist-angle \u03b8BD n Z for fixed relative limb angles of | \u03b8A n Z \u2212 \u03b8C n Z | = 90 \u25e6. However Eq. (38) may be extended to include all relative limb angles | \u03b8A n Z \u2212 \u03b8C n Z | \u2264 90 \u25e6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000029_s12239-019-0013-z-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000029_s12239-019-0013-z-Figure7-1.png", + "caption": "Figure 7. Double wishbone suspension mechanism in steering.", + "texts": [ + " In the case of the 5-SS multi-link type suspension mechanism, however, the steering axis cannot be determined geometrically as in the above cases. Kinematically, the steering axis can be defined as the instant screw axis of the wheel hub with respect to the vehicle body during its steering motion. When the suspension is steered, the wrench W3 acting along the tie rod, c0c1, in Figure 4 to Figure 6 disappears because c0, the spherical joint of the tie rod on the vehicle body, is moved by other link as shown in Figure 7, and there are only four wrenches acting on the wheel hub. The reciprocal screws to these four wrenches form a second-order screw system. This reflects the fact that the front suspension mechanism has two degrees of freedom. In order to find the steering axis for the steering motion only, it can be assumed that the wheel center remains the same height in vertical direction during the steering motion (Lee and Ahn, 1993). For this purpose, an imaginary S-PL link shown in Figure 7 that connects the wheel center, h1, and the ground can be assumed, which only constraints the vertical translation of the wheel center. Then, a zero-pitch force wrench W3 is now acting on the wheel hub through the center of the S joint normal to the ground as shown in Figure 7. With this assumption, five wrenches acing on the wheel hub is determined. Applying the same assumption for the other suspension types, the twist reciprocal to these five independent wrenches acting on the wheel determines the steering axis and its pitch for the steering motion. In this paper, the only assumption used in analyzing the steering axis is that the wheel center is acted upon by a vertical wrench that constraints its vertical motion. However, a different assumption to find the steering axis may be used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001658_icems.2019.8921632-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001658_icems.2019.8921632-Figure1-1.png", + "caption": "Fig. 1. Topologies of SM-MG and ST-MG: (a) SM-MG, (b) ST-MG.", + "texts": [ + " Then, by using the proposed HMM, a PI control of a servo system is constructed in Simulink to evaluate the dynamic performances of MGs. II. STRUCTURE CONFIGURATION AND OPERATING PRINCIPLE MGs are composed of three concentric components, any one component can remain stationary; and the other two can rotate to transmit power. The innermost and outermost components have permanent magnets. The middle component, also called as the modulator, is composed of iron bars that are evenly distributed on the circumferential direction. The two MGs studied in this paper are SM-MG and ST-MG, as shown in Fig.1 . In this paper, the outermost component is set stationary, while the other two components rotate, which are called inner rotor and outer rotor, respectively. MGs must satisfy a few structure requirements to produce a steady torque. First, the pole-pair number of inner rotor and outer rotor Pi and Po and the number of modulator pieces N should satisfy [6]: i oP P N+ = (1) 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE Besides, Pi, Po and N should be relatively prime to one another. Then, to produce a steady output torque, the rotating speed of inner rotor and outer rotor \u2126i and \u2126o must obey [14]: i 0i oP N\u03a9 \u2212 \u03a9 = (2) Consequently, based on the law of conservation of energy, the torque ratio of inner rotor and outer rotor can be expressed by [15]: i o i T N G T P = = (3) III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002354_b978-0-12-818471-4.00006-6-Figure6.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002354_b978-0-12-818471-4.00006-6-Figure6.1-1.png", + "caption": "Figure 6.1 Schematic definition of 4D printing. Terms in blue text highlights and differentiates 4D printing enhancements from 3D printing. Adapted from F. Momeni, S.M.M. Hassani N, X. Liu, J. Ni, A review of 4D printing, Mater. Des. 122 (2017) 42e79. Reproduced with permission from Elsevier \u00a9 2017.", + "texts": [ + " In order to attain dynamic capabilities, apt selection of smart materials occupying a 3D space is required [9]. The application of mathematical modeling is necessitated for ensuring that the multiple materials incorporated are appropriately designed and distributed. The 4D structure should possess a minimum of two stable states that shift from one to the other in the presence of an appropriate stimulus [10]. A schematic depiction highlighting the progression from 3D printing to 4D printing is provided in Fig. 6.1. Advanced 3D-Printed Systems and Nanosystems for Drug Delivery and Tissue Engineering https://doi.org/10.1016/B978-0-12-818471-4.00006-6 Copyright \u00a9 2020 Elsevier Inc. All rights reserved. For a 4D printing facility, the following components are integral: - A 3D printing facilityd3D printing is employed to build the 4D-printed structure from multiple materials with the desired distribution and architecture as a single one-time printed structure (i.e., multimaterial structure with simple geometry)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-Figure7-1.png", + "caption": "Fig. 7. Linear motion experiment device", + "texts": [ + " The gear G2 is attached to the shaft passing through the interior of the square shaft S1, S2, and the shaft is fixed to the square shaft S1, so the rotation of the gear G2 transmit to the square shaft S1. Then, the rotation of the square shaft S1 with respect to the square shaft S2 is demonstrated. The motor M3 drives the gear G3, and the entire internal structure rotates around the internal gear G4. Then, rotate the square shaft S1, S2 with respect to the base. We performed an operation experiment of the linear mechanism. Fig. 7 shows the overview of experiment device. The elastic telescopic structure has a total length of 3440 mm for 10 nodes. In order to control, the command of the rotation angle of the motor was sent from the PC to the microcomputer by serial communication, and the microcomputer calculated the speed command value of the motor and sent the signal to the motor driver. In addition, the ropes were attached the tip of the 1st, 4th and 7th nodes and pulled by hand to generate the tensions of the ropes used for contraction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002331_humanoids43949.2019.9034995-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002331_humanoids43949.2019.9034995-Figure3-1.png", + "caption": "Fig. 3. Sliding mass system used to replicate the mass distribution of TORO.", + "texts": [ + " It features 27 joints (excluding the hands), which are located in the legs (2x6 DoF), the hip (1 DoF), the arms (2x6 DoF), and the neck (2 DoF). The mockup was used to conduct the fall tests, since TORO was not designed to sustain the impact of a fall [13]. The dummy is constructed from standardized aluminum profiles, recreating the same kinematic structure as the actual robot. To match the weight and mass distribution of TORO, the links of the dummy are equipped with additional weights made of steel (Fig. 3). The mockup is articulated, to facilitate the change of postures for different test conditions. During a test, all joints were mechanically locked in order to preserve the desired pose. The tests were carried out with and without airbags, and using protective gear (pads for elbows and knees used in skating and skateboarding) to minimize damages on the mockup and the floor. For protecting the torso and backpack, a commercial inflatable airbag jacket by HELITE (Fig. 4) was selected; it is worn by motorcyclists to protect them in case of an accident" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002846_012072-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002846_012072-Figure2-1.png", + "caption": "Figure 2. The overall view of experiment set-up that can be used to measure tire characteristics. LVDT 1 and LVDT 2 are used to measure the contact patch length and the plate displacement, respectively. Load cell 1 is used to measure normal force.", + "texts": [ + " A braking efficiency can be calculated by dividing the maximum braking force with normal force. With this device, it is possible to measure not only braking efficiency, but also tire characteristics. This can be done if braking force resulted from vehicle braking system is strong enough to make the wheel in a locked condition. In other words, the wheel does not rotate while the plate is moving. However, more sensors will be needed to measure parameters representing tire characteristics. The experiment set-up design to do a tire test according to the concept in figure 1 is shown in figure 2, 3, 4, and 5. It requires two load cells and two LVDT (Linear Variable Differential Transformer). Load cell 1 is used to measure normal force while load cell 2 is used to measure friction force. LVDT 1 is used to measure the contact patch and LVDT 2 is used to measure plate displacement. The main plate ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10.1088/1757-899X/788/1/012072 is moved manually from the input wheel. Power from input wheel is transmitted to the main plate through gear box, ball screw, secondary plate, and load cell 2", + " Therefore, the distance travelled by the wheel will be further compared to the wheel with no deformation . This effect also can be explained by figure (b). Deformation S makes the effective radius (re) longer than the free rolling radius (r0). As a result, with the same angular velocity, the braked wheel will travel further than the non-braked wheel. Slip can be defined as the percentage of distance difference travelled by a wheel. Its difference is calculated between braking and free rolling conditions. If this definition is used in the experiment setup design as in figure 2, slip can also be defined as percentage deformation to the initial length. Therefore, the relationship between slip and the plate displacement can be written as equation (3). (3) where i is slip, S is the total deformation, S0 is the initial length, x is the plate displacement, lt is the length of contact patch, is part of the tire tread in front of contact patch that stretches. 2.4. Velocity of the main plate According to figure 2-5, the velocity of the main plate can be calculated with equation (4). (4) where Vp is the main plate velocity, nin is an angular velocity of the input wheel, zg is a gear ratio of the gear box, zb is a gear ratio of the ball screw. ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10.1088/1757-899X/788/1/012072 2.5. The algorithm of data analysis program The aim of data analysis is shown in figure 10. Data analysis is used to process input data to be output data. Input data are obtained from a tire test using the brake tester as in figure 2. Those inputs are friction force against plate displacement, length of contact patch, and normal force. Output data are friction coefficient, tangential stiffness, slip at maximum friction force, and length of . With those output data, tire characteristics can be modelled using Julien\u2019s Theory. There are some terms in Julien\u2019s Theory. These terms can be seen in figure 11 showing the relationship between friction force and the slip of a tire. Julien\u2019s Theory only models from point O to B. There are two regions between point O and B", + " If that condition is fulfilled, the program ends. The output data in the form of friction coefficient, , tangential stiffness, and the slip at maximum friction force can be displayed. With those output data, tire characteristics can be modelled using Julien\u2019s Theory. A program made with MATLAB Script for analysing the input data based on figure 12 has been made. This program needs a verification to check its correctness. To do that, the plate displacement and friction force are simulated according to the brake tester design shown in figure 2. Those plate displacement and friction force data are used as an input of the program. After that, the results of the program are compared to the actual value obtained from a reference. ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10.1088/1757-899X/788/1/012072 A simulation of friction force and plate displacement is done based on tire characteristics data from a reference [28]. These tire data are as follow: = 31.5 mm lt = 282 mm p = 0.888 kt = 4610304.1 N/m2 W = 4000 N From the tire data above, the slip at maximum friction force (im) can be calculated with Julien\u2019s Theory as in equation (11). (11) The data of input wheel rotation (nin,), the ratio of the gear box (zg), and the ratio of the ball screw (zb) used in the experiment design as in figure 2 are: nin = 1 revolution/second zg = 1 : 60 zb = 5 mm/revolution Using equation (4) and data above, the velocity of the plate is: Vp = 1 / 60 x 5 mm/s = 0.0833 mm/s (12) The data of the friction force versus the slip can be calculated with equation (8) and (9). Then, those slip data are converted to plate displacement with equation (3). Plate displacement data can be converted to time-based data with equation (12). Hence, the data of friction force versus slip, plate displacement, and time can be plotted as shown in figure 13. Slip, plate displacement, and time needed from zero to the maximum friction force are shown in figure 13. Peak point is achieved when the slip is around 9%. Plate displacement needed to reach its peak point is about 27 mm. It needs around 350 seconds to take data from the plate, which is in rest until moving at distance around 27 mm. There is enough time to observe the phenomenon of friction force using the experiment set up as in figure 2. Only friction force versus plate displacement is used as the input for the data analysis program based on the flow chart in figure 12. 3.2. Results of data analysis program When the friction force versus plate displacement data in figure 13 are used as the input for the MATLAB Script based on figure 12, the results are the following tire characteristics: = 31.5 mm kt = 4610304.1 N/m2 p = 0.888 im = 8.6871 ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003693_icra40945.2020.9196516-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003693_icra40945.2020.9196516-Figure8-1.png", + "caption": "Fig. 8: Pendulum trajectories and corresponding light painted letter shapes. Each highlighted segment on the right consists of a set of waypoints traversed in quick succession. Note that the complete trajectories may be quite long, as seen in the traces on the left, and it is virtually impossible to design such trajectories by hand or using kinematic path planning.", + "texts": [ + " In addition, a software limit is imposed on the input voltage signal u to avoid damaging the motor. To account for the joint and control limits, the following inequality constraints \u2212umax \u2264 ut \u2264 umax, \u2212\u03b8max \u2264 \u03b8t \u2264 \u03b8max are added to our direct collocation formulation of the trajectory optimization problem described in Section III. For all of our experiments, the initial state x0 is assumed to be zero, which corresponds to the system being still, with the pole centered in the front and hanging down. Following the pipeline from Fig. 4, the trajectories shown in Fig. 8 were obtained in simulation. The traces on the left show that a significant amount of time is spent in preparation of each maneuver, while the pendulum is accumulating the required energy and momentum to pass through the waypoints in the specified order and in quick succession. The visualizations on the right show the expected results from the light painting photography, where the letter segments are highlighted based on the activation times t\u0302i obtained through the \u2018attention\u2019-augmented trajectory optimization described in Section III-C", + " Long exposure photographs of the light painted letters are presented in Fig. 9. The pictures have been taken in a dark room with an LED device synchronized with the trajectory execution and activated based on the optimized 1509 Authorized licensed use limited to: La Trobe University. Downloaded on September 21,2020 at 13:06:09 UTC from IEEE Xplore. Restrictions apply. segment beginning/end times t\u0302i described in Section III-C. Comparing the real images in Fig. 9 with the simulated renderings in Fig. 8, we observe a sufficiently good match allowing the letters to be well recognizable. However, the trajectories slightly deviate towards the end, as it can be seen on the middle strokes in the letters \u2018A\u2019 and \u2018S\u2019 that are drawn last. These segments are slightly tilted compared to their desired location. For a better view, see the accompanying video, where real and simulated trajectories are drawn side by side. A method for objective function design in the context of trajectory optimization with waypoints has been presented (see Section III)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000477_978-3-030-20131-9_72-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000477_978-3-030-20131-9_72-Figure1-1.png", + "caption": "Fig. 1. The therapeutic needle guiding module", + "texts": [ + " From the kinematic point of view, the motions and degrees of freedom of the two robotically manipulated instruments are as follow: - the intra-operatory ultrasound (US) probe, a Hitachi Arietta UST 9150 [8], will be inserted through an incision that will act as a remote center of motion (RCM) requiring a total of 5 DoF from the robotic device that will ensure the 3 DoF for its positioning after the insertion (as the RCM behaves as a class 2 joint with four DoF that has to be maintained fixed within the robot workspace) to which the module that holds the US probe will have an additional of 3 DoF allowing the rotation of the probe around its longitudinal axis and the bending of its distal head using the two build-in handles; - the therapy needle will require 5 DoF being inserted in a two steps motion, whereas based on the definition of the insertion and target points, the needle will be positioned on the predefined trajectory outside the body (a 5 DoF motion) and then inserted on a linear trajectory (1 DoF motion) similar to the approach described in [9]. Based on the data presented above, Section 2 will describe the concept of a new modular parallel robotic system with two independent robotic arms, followed by its kinematic model, singularities analysis and workspace modelling. Based on the motion requirements of the two instruments that will be manipulated using the robotic device, a system, PRoHep-LCT [10], having two independent parallel modular arms is proposed. The first arm, used for the positioning of the therapeutic needle (see figure 1) consists in a parallel module with M=3 DoF of family F=1 and three active joints ( 1 2 3, ,q q q ) that generate a spatial motion with constant platform orientation and a second module working in cylindrical coordinates having M=3 DoF of family F=1 and two active joints ( 4 5,q q ). The two modules are connected through a pair of Cardan joints with the needle module mounted between them. Referring to the first module, the active joint 1q performs a translation along the Z axis, while the active joints 2q and 3q perform rotations around axes parallel with the Z axis", + " (6) provides the coordinates of the point 1A : 1 1 1 1 1 2 , 1,2A i R i A i R i X X c i Y Y c (7) The Z coordinate of the point 1A is: 1 1 1A qZ q d (8) The coordinates of the second Cardan joint are determined as follows: 2 2 2 5 4 5 4 4 min 2 4 2 2 2 1 1 2 ( ) ( ) , 1,2 A q q q us A A i A i p A i A i X l q q l l Y q Z Z l X X i (9) The orientation angles of the needle mobile platform are: 1 2 1 2 arcsin , 1,2 arcsin , A i A i p A i A i p X X l i Y Y l (10) The coordinates of the end-effector are determined with the following equation: 1 1 1 cos( ) sin( ) sin( ) sin( ) , 1, 2 cos( ) Ei A i Ei A i E A X X d Y Y d i Z Z d (11) The geometric modelling for the second parallel module, used of the positioning of the ultrasound probe has similar equations as its kinematic structure is identical and will not be detailed further on. The workspace of the parallel module has been assessed initially by computing the total operational workspace using the equations of the inverse geometrical model, for the following set of geometrical parameters, expressed in [mm] (see figure 1): 1 2 3 1 2 4 5 4 min 120; 115; dq 140; 315; 530; 150; 80; 75; 100; 325; 400; 25; p linkd l d lq lq c c dc lq lq lq (12) The resultant workspace is presented in figure 4. The total workspace representation shows relevant data regarding the volume that the robot can cover but almost no information about the needle orientation capability. Due to the fact that, for each procedure the robot will be positioned relative to the patient in a convenient way (based on the pre-operative data) the authors conducted an analysis of several points located in the central area of the robot workspace to evaluate its orientation capability" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001661_icems.2019.8921927-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001661_icems.2019.8921927-Figure6-1.png", + "caption": "Fig. 6. Rotor mechanical stress nephogram. (a) non-uniform. (b) uniform.", + "texts": [ + " 3) The rotor core is assumed as an entity and the lamination effect on material properties is negligible. The International Electro Technical Commission standard IEC60034-1 stipulates that the motor should be able to withstand 1.2 times the maximum speed. Therefore, the mechanical state under the rotor with a rotational speed of 14400rmp is taken as the research object when performing stress simulation. The mechanical stress simulation nephograms and total deformation nephogram of the two rotors are shown in Fig. 6 and Fig. 7. The simulation results show that the maximum mechanical stress (MMS) of rotor with uniform air gap is 456.68 MPa, a little smaller than the rotor with non-uniform air gap , which is 464.27 MPa. The MMS point of both rotors on bilateral bridge. In this paper, the structure of the outer ring of the rotor does not affect the position of the MMS point and the MMS value. Moreover, the total deformation of the two rotors is basically the same. IV. CONCLUSION The simulation results of one particular design are demonstrated in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001237_s11517-019-02044-4-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001237_s11517-019-02044-4-Figure5-1.png", + "caption": "Fig. 5 Direction of axis of rotation and angle of rotation around axis", + "texts": [ + " The combination of the information from these three sensors enables the estimation of the real orientation of the object in 3D space [36]. The orientations of the PS Move have been saved as the quaternion method [41] in a laptop during the 3DUS reconstruction. A quaternion is a mathematical concept that represents the orientations and rotations of an object in 3D space. Quaternion is used as a descriptive axis direction of rotation and total angle of rotation around the axis [42], as shown in Fig. 5. Direction of an axis of rotation and total angle of rotation is represented in Eqs. 1 to 4: Angle = cos(q/2) (1) Axisx = qx/ \u221a 1 \u2212 q2 (2) Axisy = qy/ \u221a 1 \u2212 q2 (3) Axisz = qz/ \u221a 1 \u2212 q2 (4) where the total angle of rotation is known as Angle. The axes of rotation in x-direction, y-direction, and z-direction is denoted as Axisx, Axisy, and Axisz respectively. The quaternion values are represented as qx, qy, and qz. Due to its numerically more stable and efficient characteristics, the quaternion method was chosen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure10-1.png", + "caption": "Figure 10. Stephenson2 mechanism (2).", + "texts": [], + "surrounding_texts": [ + "The primary objective of this study was to identify the joint that producesthe highest error in a mechanism due to clearance. Simulations were carried out for all mechanisms described in the previous section by introducing clearances in different joints, one at a time. The effect of different speeds (of crank) and clearance sizes was also studied. Some important results are discussed here. The results for the four bar mechanisms have been summarized in figures 13, 14 and 15. These show that Joint 3 is the most critical joint and the values vary significantly for 0.02 mm but they are very similar for 0.5 mm. So, sensitivity to speed is high for low clearance. Figures 16, 17 and 18 show mean deviation at different joints of different inversions of the Watt chain at 0.1 mm radial clearance. The trend of all the curves is the same. Joint 3 is the most critical joint in all the three mechanisms followed by joints 2, 4 and 5, i.e., the criticality order here is: Joint 3[ Joint 2[Joint 4[ Joint 5. Joint 6 in figures 5 and 6 is the grounded joint which is least critical of them all. Figures 19 and 20 show the criticality graphs for Stephenson1 mechanism and Stephenson1 mechanism with slider, respectively. Both the graphs show the same trend where the deviation is maximum at the output link and decreases as one proceeds to the other end of the mechanism. A common trend which seems to emerge is that the first joint next to the crank, which has a link connected to ground, is the most critical joint. After that the criticality ranking goes towards the crank and then goes further down to the joints in order of their distances from the crank. We would examine the validity of this hypothesis in the rest of the study. Figures 21 and 22 show the criticality graphs for Stephenson2 mechanism (1) and Stephenson2 mechanism (2). The order of criticality for both the mechanisms is: Joint4 [ Joint3 [ Joint2 [ Joint5. This is as per the hypothesis stated earlier. The most critical joint is the joint next to the crank, which has a link connected to the ground (i.e., Joint 4). The next most critical joints are the ones towards the crank, i.e., Joint 3 and Joint 2. Then comes the joints which are left, i.e., Joint 5. Joint 6 is the grounded joint. The two graphs look very similar as the structures of mechanisms are the same. The deviations increase significantly with increase in speed, if the clearance is at Joint 4. The values in first graph are slightly higher than the corresponding values in the second graph. The overall conclusion is that deviations are more, if the output link is closer to the crank. Similar results are seen in eight and ten-bar mechanisms as shown in figures 23 and 24. The joint next to the crank is the most critical joint. It should also be noted that the grounded joints also show a trend. The grounded joint in the first loop has a significant effect but rest of the grounded joints have very little effect on the output. Error in output generally increases with increase in the crank speed (expect for an aberration shown in figure 19). This happens for all clearance sizes. However, the change in mean deviation is quite small even with a 5 times increase in speed from 100 to 500 rpm. Figures 25 and 26 show the variation of mean deviation with increase in speed when a clearance of 0.1 mm is at Joint 2 and Joint 3, respectively. It can be noticed that the deviations usually increase with speed. This means that a machine working at higher speeds is expected to produce more error than a machine working at lower speeds. Increase in radial clearance size brings an almost linear increase in the deviation values as seen in figures 27 and 28. It is also worth noting from figure 27 here that the deviation of slider is nearly 50% for 0.1 mm clearance (0.0501 is 50% of 0.1) and it goes on increasing to 80% for 0.5 mm clearance (0.3977 is 80% of 0.5). This happens because at lower values of clearance, the number of collisions between journal and bearing is more. Due to more number of collisions, the journal is sent back to the center again and again, and thus it remains at the center for a larger amount of time as compared to the case of large clearances. Thus, in larger clearances, the journal appears to remain closer to the walls of the bearing for an extended period and spends less time going to the diametrically opposite wall." + ] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-Figure2-1.png", + "caption": "Fig. 2. Concept of using tendon-driven elastic telescopic arm (a) Observation of remote areas, and (b) Improvement of stepping ability by mounting on a mobile robot", + "texts": [ + " Furthermore, since the arm is long, the range of movement at the tip position of the arm can be large by rotating the base itself. On the other hand, since the displacement of the tip position due to bending does not depend on the arm length but the rope winding length, it is possible to control the tip position accurately with a small reel diameter. The proposed arm can be more compact than conventional ones, so the restriction on the installation location is reduced. For the reason it is suitable for using in a disaster site as shown in Fig. 2a and for mounting on a mobile robot as shown in Fig. 2b. On the other hand, since it is composed of a thin elastic body, a large structural vibration is easily expected. However, it may be possible to suppress vibration if it is combined with bending control by tendon drive. In this section, we propose a mechanism to achieve the linear motion of the arm. In addition, we show the result of the linear motion experiment by using a manufactured mechanism. As for extending motion of the elastic telescopic structure, we proposed a method of pushing out from inside it using a stand tube [11]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003750_physreve.102.033115-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003750_physreve.102.033115-Figure1-1.png", + "caption": "FIG. 1. A singly flagellated bacterium swims straight by rotating its flagellum straight behind the cell body, and reorients by a flick of its flagellum during which the flagellum undergoes large deflections (characterized by \u03b8 ) associated with large nonlinear deformations of the hook.", + "texts": [ + "edu backward runs by changing the rotation of its single flagellum at one pole from counterclockwise, with the flagellum pushing the cell, to clockwise, with the flagellum pulling the cell. During clockwise rotation of the flagellum, the hook unwinds and its stiffness decreases. Upon switching to counterclockwise rotation, the pushing flagellum places the unwound hook under compression and, after a short time, dynamical buckling instabilities of the hook cause the flagellum to make large deflections from the cell-body axis (a flick; Fig. 1) and reorient the entire bacterium [4,5,13]. During this process, the counterclockwise rotation winds the hook and increases its stiffness, and eventually it becomes stiff enough to stop buckling. Then, the bacteria runs forward in a straight trajectory with its flagellum on-axis. Thus, during preflick runs, the hook is unstable to large deflections, while during postflick runs, the hook is stable to large deflections. To investigate the role of hook flexibility in different aspects of bacterial swimming dynamics, a number of different models have been used in previous studies", + " In our previous work [5] investigating flick motility, we showed that the hook also has an important torsional response, and modeled it as a combination of linear bending and torsional springs with stiffnesses calculated from the linearization of a Kirchhoff rod model. We showed that this simple model for hook bending, combined with a model for swimming dynamics that ignores hydrodynamic interactions between the flagellum, hook, and cell body, can predict the experimentally observed torques needed to initiate flicks. It seems likely that for large bending, complex threedimensional (3D) deformations (such as shown in Fig. 1) may become important. While simple linear spring models allow efficient computation and investigation for a wide range of cell body, flagellum, and hook parameters, linear models of hook response are not adequate for large hook deformations. Instead, Kirchoff rod models of the hook can be used to describe large hook deformations. Shum et al. investigated the dynamics of free-swimming singly flagellated bacterium, including the instability of straight swimming, by combining a Kirchoff rod model of the hook with a boundary element method to treat hydrodynamic interactions [17]", + " We do this by studying progressively more faithful models for the hook, starting with the previously employed (1) linear bending springs and (2) linear bending and torsional springs, then (3) models that incorporate nonlinearities due to larger hook deformations, (4) Kirchhoff rod models that fully incorporate large hook deformations, and, finally, a (5) numerical rod model that also incorporates the hydrodynamic interactions of the hook. We also employ models (1) through (4) both with and without hydrodynamic interactions between the flagellum and cell body to test the importance of those hydrodynamic interactions. We show that for stiff hooks, a bacterium swims with its flagellum rotating on-axis. As the hook stiffness decreases, there is a transition to swimming in which the flagellum precesses about the axis with a large deflection angle \u03b8 (Fig. 1). The critical hook stiffness at which the transition occurs depends on the cell-body and flagellum geometry, hook model, and hydrodynamic interactions. For all hook stiffnesses below the critical value, linear spring models predict precession; however, for hook stiffnesses only slightly below the critical value, nonlinear effects destabilize precession into complex dynamics with large hook deformations. For these unstable deflections, we find that the choice of nonlinear models or Kirchoff rod models, as well as the inclusion of hydrodynamic interactions, can produce significantly different flagellum and cell-body trajectories" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003853_ecai50035.2020.9223138-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003853_ecai50035.2020.9223138-Figure9-1.png", + "caption": "Fig. 9. Isometric view of the final device", + "texts": [], + "surrounding_texts": [ + "The electronic design consisted in the realization of the block diagram of the device and in the calculation of the electronic control parameters in accordance with the nominal operating values of the decontamination device [10-14]. The main elements of the electronic control unit are: an accumulator with the role of providing the electricity necessary for the operation of the entire device, a speed variator for a DC motor (DC-DC converter), an LCD display and, optionally, a UV lamp used to sterilize the HEPA filter. The ventilator chosen to supply the decontaminated air should normally allow about 1/2 liter of air to enter the respirator during respiration and is eliminated in the same amount. Considering that a healthy adult breathes 16 times a minute, it can be calculated that in an hour about 480 liters of air are carried through the lungs. From this air, an adult consumes about 300 cm3 of oxygen per minute, which is 18 liters of oxygen in an hour. The chosen fan, model Sunon MAglev MEC0381V1, 120mm, shown in the figure 11 offers an air flow between 12 and 75 cubic meters / minute . This is equivalent to an air flow of 720-4500 liter/hour. Fabricantul specifica o durata de functionare de 70.000 ore. The power consumption of the fan is 10W / The electrical efficiency of the converter of a dial used in the control of the fan speed is about 98%. DC motor speed variator To change the fan speed and implicitly to change the air flow we use a DC-DC converter [15-19]. The electrical characteristics of this converter are the following: - supply voltage: 4.5 - 35V d.c. - maximum power of the controlled load: 90W; - maximum current by load (DC motor): 2.6A Fig. 13. DC-DC converter for sped variator Figure 13 shows the electrical diagram of the converter, and Figure 14 shows the result of simulating the operation of the speed controller. The electric motor is modeled Rm-Lm. In figure 14, the waveforms represent the voltage on the motor depending on the duty cycle of the received PWM command. In average values, the voltages on the motor are: 9V - green, for filling factor 50%; 10.5V \u2013 red, for 75% fill factor, 11.4V \u2013 blue, 90% fill factor. Duty cycle [%] DC motor voltage [V] Airflow [liter/hour] Power consumption [W] 50% 9 3375 7,5 75% 10,5 3937,5 8,75 90% 11,4 4275 9,5 Battery Under the conditions imposed by the electrical design, the maximum total consumption of the device is 11W / h. To obtain an operating autonomy of at least 8 hours (a normal program of a doctor) we need a power source that provides us with 88W / cycle. From the previous table and taking into account the required air flow, an energy consumption of 5W / h results. To ensure this energy consumption we need a battery with approximately 4000mAh. The chosen battery, ABENIC DC 12V 10A (120W) 9800mAh Super Rechargeable Protable Li-ion Lithium Battery DC1298A, corresponds to our electricity requirement and provides us with 9800 mAh. At the same time its weight is very small \u2248 250g. Authorized licensed use limited to: Middlesex University. Downloaded on November 01,2020 at 15:38:55 UTC from IEEE Xplore. Restrictions apply. LCD display indicator device The LCD display indicator gives us information about the battery voltage level, both in numerical format and in percentage format. In order to know exactly the charge level of the battery, the possibility of a pre-programming was provided. In this phase we can enter the nominal value of the measured battery voltage [19-22]. Also, by adding a real time clock to the electrical circuit, the time of the HEPA filter was used can be displayed on the same display. The power supply of the indicator module is made at a minimum voltage of 8V. The electrical voltage of the monitored batteries must be in the range of 8-60V. The consumption of the digital display is 10mW. V. CONCLUSION The presented device supports the actions of protection of the medical personnel located in the areas with high risk of bacterial contamination. By choosing a simple electronic solution, with common components and easy to purchase in any economic conditions, but a solution that offers the possibility to control all parameters and ease of handling, the aim was to obtain a low cost. Advantages of the proposed device are: \u2022 Replace all these separate parts with an integral protective suit: mask, visor, disposable used in medical activity in hospitals. \u2022 Can be easily sterilized and decontaminated; \u2022 By minor modifications it can fulfill the function of a nebulizer for the administration of the inhalation treatment, with the control of the flow and the volume of inspired air. Further development of the device can be achieved by incorporating UV ionization lamps. This can increase the duration of use of the filter and could cause an internal predecontamination of the device. If we want to use a UV lamp, its operation should be done only when the device is not used by medical staff. In order not to affect the autonomy of operation, we recommend using the UV lamp only while charging the battery. REFERENCES [1] https://www.reginamaria.ro/articole-medicale/ne-protejeaza-masca- medicala-de-gripa-si-coronavirus [2] Bann, Darrin V.; Patel, Vijay A.; Saadi, Robert; et al. , Impact of coronavirus (COVID-19) on otolaryngologic surgery: Brief commentary, HEAD AND NECK-JOURNAL FOR THE SCIENCES AND SPECIALTIES OF THE HEAD AND NECK, DOI: 10.1002/hed.26162, 2020 [3] Howard, Rex A.; Lathrop, George W.; Powell, Nathaniel, Sterile field contamination from powered air-purifying respirators (PAPRs) versus contamination from surgical masks, AMERICAN JOURNAL OF INFECTION CONTROL, Volume: 48, Issue: 2, Pages: 153-156, DOI: 10.1016/j.ajic.2019.08.009 [4] Xu, Susan S.; Lei, Zhipeng; Zhuang, Ziqing; et al. COMPUTATIONAL FLUID DYNAMICS SIMULATION OF FLOW OF EXHALED PARTICLES FROM POWERED-AIR PURIFYING RESPIRATORS, Conference: ASME International Design Engineering Technical Conferences / Computers and Information in Engineering Conference Location: Anaheim, CA Date: AUG 18-21, 2019 [5] Lee, Min-Ho; Yang, Wonseok; Chae, Nakkyu; et al., Performance assessment of HEPA filter against radioactive aerosols from metal cutting during nuclear decommissioning, NUCLEAR ENGINEERING AND TECHNOLOGY Volume: 52 Issue: 5 Pages: 1043-1050 Published: MAY 2020 [6] Dehghan, S. F.; Golbabaei, F.; Mousavi, T.; et al., Production of Nanofibers Containing Magnesium Oxide Nanoparticles for the Purpose of Bioaerosol Removal, POLLUTION Volume: 6 Issue: 1 Pages: 185-196 Published: 2020) [7] Bogue, R. (2013). 3D printing: the dawn of a new era in manufacturing?. Assembly Automation. [8] Alsoufi, M. S., & Elsayed, A. E. (2018). Surface roughness quality and dimensional accuracy\u2014a comprehensive analysis of 100% infill printed parts fabricated by a personal/desktop cost-effective FDM 3D printer. Materials Sciences and Applications, 9(1), 11-40. [9] \u0106wik\u0142a, G., Grabowik, C., Kalinowski, K., Paprocka, I., & Ociepka, P. (2017, August). The influence of printing parameters on selected mechanical properties of FDM/FFF 3D-printed parts. In IOP Conference Series: Materials Science and Engineering (Vol. 227, No. 1, p. 012033). IOP Publishing. [10] Promise Elechi. An Improved Robotic Control System Using Wireless Fidelity Network. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2020 6(2):1-6 [11] Alexandru Calin Stan, Oprea Mihaela. Petri Nets Based Coordination Mechanism for Cooperative Multi-Robot System. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2020 6(2):7-14; https://jeeeccs.net/index.php/journal/article/view/138 [12] Promise Elechi, Sunny Orike. Application of Light Fidelity Network for Improved Indoor Wireless Communication System. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2020 6(2):15-20; [13] Carmen-Silvia Oprina. Secured Transmission of Data from Environments with High Potential Risk. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2019 5(4):7-10; [14] Anshu Prakash Murdan, Luckshana Gunness. An Internet of Things based system for home automation using Web Services and Cloud Computing. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2017 3(1):29-36; [15] Paul Owoundi Etouke, L\u00e9andre Nneme Nneme, Jean Mbihi. An Optimal Control Scheme for a Class of Duty-Cycle Modulation Buck Choppers: Analog Design and Virtual Simulation. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2020 6(1):13-20; [16] Paul Owoundi Etouke, L\u00e9andre Nneme Nneme, Jean Mbihi. An Optimal Control Scheme for a Class of Duty-Cycle Modulation Buck Choppers: Analog Design and Virtual Simulation. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2020 6(1):7-12; accesed on March 9, 2020; [17] Leandre Nneme Nnem, Bertrand Moffo Lonla, Gis\u00e8le B\u00e9atrice Sonfack, Jean Mbih. Review of a Multipurpose Duty-Cycle Modulation Technology in Electrical and Electronics Engineering. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2018 4(2):9-18; accesed on March 9, 2020; [18] Mihaescu M. Applications of multiport converters. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2016 2(1):13-18; accesed on March 9, 2020; [19] Sadiq Ur Rehman, Muhammad Hasnain Raza, Ahmad Raza Khan. Delta 3D Printer: Metal Printing. Journal of Electrical Engineering, Electronics, Control and Computer Science - JEEECCS 2019 5(3):19- 24; [20] Karunakara Rai, Akhilesh Jandhyala, Shantharama C. Power Intelligence and Asset Monitoring for System. Journal of Electrical Engineering, Electronics, Control and Computer Science (JEEECCS) 2019 5(3):31-36; [21] Valeriu Manuel Ionescu, Alexandru Constantin. Application for collecting data from sensors and transmitting them through the network. Journal of Electrical Engineering, Electronics, Control and Computer Science (JEEECCS) 2019 5(2):1-4; [22] F. B\u00eerleanu, N. Bizon, \u201cPrinciples, Architectures and Challenges for Ensuring the Integrity, Internal Control and Security of Embedded Authorized licensed use limited to: Middlesex University. Downloaded on November 01,2020 at 15:38:55 UTC from IEEE Xplore. Restrictions apply. Systems\u201d, Journal of Electrical Engineering, Electronics, Control and Computer Science (JEEECCS), Volume 3, Issue 7, pages 37-45, 2017; Authorized licensed use limited to: Middlesex University. Downloaded on November 01,2020 at 15:38:55 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_80_0000333_1.5096097-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000333_1.5096097-Figure2-1.png", + "caption": "FIG. 2. Lightweight stiffener profile with functionally integrated processing aids (Ref. 10) [Reprinted from D.-A. T\u00fcrk et al., \u201cComposites part production with additive manufacturing technologies,\u201d Proc. CIRP 66, 306\u2013311 (2017)].", + "texts": [ + "6 The specific combination of different structures, materials, and/or manufacturing techniques can enable the systematic tailoring of product properties, such as stiffness, damping effects, or thermal and electrical characteristics.1,6\u20138 A comprehensive survey on the integration of different types of sensors into additively manufactured parts for a broad range of applications was published by Lehmhus et al.9 T\u00fcrk et al. presented the application of the functional integration principle on the basis of several case studies, including lightweight stiffener profiles (Fig. 2) for winglets and novel aircraft instrument panels (Fig. 3). In the case of the stiffener profile, a carbon-fiber reinforced polymer was combined with an additively manufactured core, utilizing the freedom of design for specific adaptation of the structure to local loads and the direct functional integration of positioning and fixation elements during the manufacturing.10 Figure 3 shows the highly integrated lightweight aircraft instrument panel design developed by T\u00fcrk et al. The core is constituted of additively manufactured honeycomb structures with a variety of integrated functions, ranging from positioning and fixation elements to pockets for holding different inserts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000821_ccaa.2018.8777578-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000821_ccaa.2018.8777578-Figure3-1.png", + "caption": "Fig. 3. Throttle, pitch, roll and yaw in quadcopter", + "texts": [ + " We have achieved the CG by measuring the weights of all the components which are to be fixed on the quadcopter and then integrated them on the frame symmetrical about the center. However weight distribution depends on the type of frame we use. We have used an X-shaped frame, symmetric about its center. Stability of the quadcopter always depends on the symmetry of the quadcopter. V. QUADROTOR DYNAMICS There are four basic motions that a quadcopter can hover around. They are throttle, pitch, roll and yaw as shown in Fig. 3. Before going to the motion of the quad, we should know about the propeller dynamics. One of the diagonally related propellers should be of clockwise direction while the other propellers should be of counter clockwise direction as shown in Fig. 1. The red or the short arrow and the blue or the long arrow represents the low and the high speeds of rotation respectively. For every motion the Center of Thrust (CT) changes depending on the direction of motion of the quadcopter. For each revolution the quad moves a distance equal to the pitch of the propeller provided there is required thrust" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000881_ecc.2019.8795797-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000881_ecc.2019.8795797-Figure2-1.png", + "caption": "Fig. 2. Application of the parallel guidance law with equally fast agents.", + "texts": [ + " This is a decentralized strategy; each pursuer Pi chooses from two cases, based on angle \u03b2i (the angle \u03b2 corresponding to the ith pursuer). 1) If the evader moves such that |\u03b2i| \u2264 \u03c0/2 holds, pursuer Pi mirrors the movement, heading towards the intercept point. (Note that is in accordance with the previous section as \u03b2\u2217 = \u03c0/2 for equal speeds (5).) \u03b1i is calculated based on (3), which, with vp = ve simplifies to: \u03b1i = sin\u22121 ( sin\u03b2i ve vp ) = \u03b2i (7) The control can be given by upi = \u2220 \u2212\u2212\u2192 PiE + \u03b1i (8) 2) If |\u03b2i| > \u03c0/2, the pursuer moves parallel to the evader: upi = ue (9) Fig. 2 illustrates how both actions keep the angles \u2220 \u2212\u2212\u2192 EPi. Note that this strategy is applicable only when the pursuers have an informational advantage, i. e. know instantly what control action the evader chooses. In practical implementations, the pursuers often lack this information and have to approximate. Extensions to the superior evader case where \u03b2\u2217 < \u03c0/2 exist. The reference strategy proposed by Ahweda and Schwartz [1] will be denoted by \u0393p. Note that this strategy is independent of the applied robot model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002512_icase48783.2019.9059099-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002512_icase48783.2019.9059099-Figure6-1.png", + "caption": "Fig 6: Shows the (a) airfoil shroud and (b) mounting internal arrangement used by Passmore [4]", + "texts": [ + " Another advantage of decreasing the shaft length was that the ball was now positioned at equal distance from the tunnel floor and tunnel roof after its attachment with the balance. Another step taken was using an airfoil and attach it with the support shaft. The Airfoil used was NACA Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 14,2020 at 08:37:27 UTC from IEEE Xplore. Restrictions apply. 978-1-7281-1955-7/19/$31.00 \u00a92019 IEEE 0021 and was made out of model board. The setup is shown in figure 6. Further vibrations were reduced by addition of a shroud, this allowed measurements at airflow speeds of 30 m/s. A constant frequency (300 Hertz) was used to measure and collect data from spin balance. This data was then processed utilizing a binning method. In binning method, a bin is assigned to every data point corresponding to its angular orientations. Each data point set included averaged readings of twenty rotations of the ball. Taking the average of the readings facilitated in determining the isolated effects of ball orientations free from the inherent non steady facotrs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure90.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure90.2-1.png", + "caption": "Fig. 90.2 Representation of plug, fillet and CMT welding locations in the liner", + "texts": [ + "1 Physical properties of AA 6061-T6 Young\u2019s modulus, MPa Poisson\u2019s ratio Density, kg/mm3 73,100 0.34 2.76 10\u22126 Table 90.2 Mechanical properties for base and weld metal AA 6061-T6 90 Finite Element Analysis of Potential Liner Failures \u2026 1075 sheet using die and punch system, and forged part for end boss is obtained by machining a solid block of AA 6061-T6. The parts are further joined by plug welding technique, and welding the hemispheres circumferentially using cold metal transfer welding (CMT) technique as shown in Fig. 90.2. The true distance between the points of laying plug weld to edge of fillet weld is maintained at 8 mm with each plug having radius of 5 mm. Between each plug weld, a distance of 21 mm is maintained and positioned at three stages with pitch angle of 20\u00b0 as shown in Fig. 90.2 (section A\u2032\u2013A\u2032). The liner is modelled as solid part. The effect of welding process changes the inherent base material properties in aluminium. Hence, by considering the effects due to welding, weld metal properties are assigned at pole and equator region, and length to be assigned is determined from an experimental trial. The length for assigning weld metal properties in liner is determined from Fig. 90.3. A miniature level AA 6061-T6 tube test specimen (replicating the large-scale liner structure) having a thickness of 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002087_012059-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002087_012059-Figure4-1.png", + "caption": "Figure 4. Pressure Angle of Cycloidal Tooth Profile", + "texts": [ + " From figures 2 and 3,it can be seen that there are inflection points in the middle of the ICMAE 2019 IOP Conf. Series: Materials Science and Engineering 751 (2020) 012059 IOP Publishing doi:10.1088/1757-899X/751/1/012059 tooth side. The curvature of cycloid tooth profile changes quickly near the tooth root, and the curvature change near the tooth top is smaller than that near the tooth root.In the convex tooth profile section,the minimum curvature is not at the tooth top,but near the tooth top. Pressure angle is an important index for judging transmission performance. From figure 4,In order to describe the transmission characteristics between cycloid gear and pin teeth,the angle between the relative speed of rotation of cycloid profile relative to geometric center and the force direction of cycloid profile at meshing point is defined as the pressure angle between cycloid profile and pin teeth. By definition,the expression of pressure angle can be describe as follows: (11) In the equation (11), is the force acting on the cycloid tooth profile at the meshing point K,The direction is the same as the normal vector of the cycloid profile at point K" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002558_j.procs.2020.03.371-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002558_j.procs.2020.03.371-Figure2-1.png", + "caption": "Fig. 2. Fragmented tooth and Missing tooth Gears", + "texts": [ + " An acceleration sensor also mounted on the gearbox to study the amplitude of the healthy and faulty gearbox. Gearbox has a gear-pinion arrangement. Gear has 27 and Pinion has 18 teeth. Fig.1 shows the Experimental setup. In the experimental investigation, the features of the raw vibration response are evaluated and studied in three dissimilar situations of gearbox i.e., healthy tooth, fragmented tooth and missing tooth gearbox under the three dissimilar speed conditions i.e., 15Hz; 25Hz; and 35Hz with three different loading conditions with the time gap of 30 Sec. Fig.2 Fragmented tooth and Missing tooth Gears. Fig.3 (a) shows the conventional method of vibration signal pre-processing which is directly dealing with TSA or EMD. The main limitations of this method are (a) Conventional Pre-processing method was unable to pre-process gearbox vibration signal because of its low noise ratio and (b) Sensors have to be mounted in the smallest part of any machinery, which can make many cases of access restriction such as sometimes vibration signals recorded far away from the gearbox and produced more intrusions due to the long path of vibration", + " An acceleration sensor also mounted on the gearbox to study the amplitude of the healthy and faulty gearbox. Gearbox has a gear-pinion arrangement. Gear has 27 and Pinion has 18 teeth. Fig.1 shows the Experimental setup. In the experimental investigation, the features of the raw vibration response are evaluated and studied in three dissimilar situations of gearbox i.e., healthy tooth, fragmented tooth and missing tooth gearbox under the three dissimilar speed conditions i.e., 15Hz; 25Hz; and 35Hz with three different loading conditions with the time gap of 30 Sec. Fig.2 Fragmented tooth and Missing tooth Gears. Fig. 1. Experimental Setup Fig. 2. Fragmented tooth and Missing tooth Gears Fig.3 (a) shows the conventional method of vibration signal pre-processing which is directly dealing with TSA or EMD. The main limitations of this method are (a) Conventional Pre-processing method was unable to pre-process gearbox vibration signal because of its low noise ratio and (b) Sensors have to be mounted in the smallest part of any machinery, which can make many cases of access restriction such as sometimes vibration signals recorded far away from the gearbox and produced more intrusions due to the long path of vibration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001079_acc.2019.8814904-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001079_acc.2019.8814904-Figure8-1.png", + "caption": "Fig. 8. (a) Schematic of the curved space mapping Chromo-Modal Dispersion (CMD) device and the functions it performs (inset). (b) Warping of spatial dispersion by a parabolic mirror. Diagrams on the right-hand side show, from top to bottom, the mapping of optical frequency into 1-D space, 2-D space, and 1-D polar coordinate space [37]. SLM: spatial light modulator; MMF: multimode fiber or waveguide.", + "texts": [ + " The strength, sign, and the delay versus frequency curvature can be widely reconfigured by adjusting the alignment of the grating and the waveguide. The CMD device provides field reconfigurable tuning of the group delay and offers a means to achieve the type of phase filters needed for engineering the spectrotemporal structure of wideband optical waveforms in a reconfigurable fashion. Also recently, linear-to-curved-space mapping is combined with the CMD to achieve a new dispersive device that offers arbitrary tuning of dispersion curvature [37] (see Fig. 8). The ROADM is a tunable wavelength-division multiplexing filter with a channel monitor and attenuator/ amplifier that can be remotely reprogrammed to change the channel access. Having a wavelength selective switch, they can be used in parallel in conjunction with a set of tunable delays to form a quantized form of the group delay profile. As long as the spectral resolution of the channels in the ROADM is fine enough for the target application, this can be an effective approach to implement an easily reconfigurable arbitrary group delay profile" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001567_j.ifacol.2019.11.260-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001567_j.ifacol.2019.11.260-Figure1-1.png", + "caption": "Fig. 1. Jetstream-3102 aircraft withe the referential frames configuration.", + "texts": [ + " These control laws also allow specific state variables to be commanded directly, and remain valid over the entire flight envelope without requiring gain scheduling. In this paper, we combine the NDI control approach with HHOSM to produce the control system for the attitude and velocity-tracking problem of a fixed wing aircraft. Thus, the main goal is to develop a reliable and wellperformed autopilot so that attitude and longitudinal speed of the aircraft track the desired attitude and desired velocity respectively under system uncertainties. The dy- Fig. 1. Jetstream-3102 aircraft withe the referential frames configuration. namic inversion linearizes dynamics of selected controlled variables by the use of nonlinear full-state feedback then linear controllers can be designed to regulate these variables with desirable closed-loop dynamics. To handle the uncertainty and nonlinearity, the desired closed-loop dynamics are given by an HHOSM controller. The validation and the evaluation of this approach of control is based on the use of aircraft Aircraft Control Systems (ACS)", + " The simulation environment and the results of implementation of this new autopilot are presented in Section 4. Finally, the conclusions are drawn in Section 5. 2. MODEL DESCRIPTION AND PROBLEM FORMULATION In what follows, a brief description of the main features of the aircraft is introduced as well as its dynamic model. Then, the control problem is formulated. 2.1 Aircraft aerodynamical model The chosen aircraft is a British Aerospace (Jetstream3102), which is a fixed wing, twin turboprop aircraft, illustrated in Fig.1. This type of aircraft has as control inputs the throttle setting command (\u03b4th), and the deflection angles of the three control surfaces: elevator (\u03b4e), ailerons (\u03b4a), and rudder (\u03b4r) (see Fig.1). We consider this aircraft due to the fact that all its aerodynamic coefficients are available in the literature (17),(18). The wing surface area S, the wingspan b, the mean aerodynamic chord c\u0304, and the mass m of the aircraft are considered as constant parameters given respectively as 280ft2, 46ft, 6.5ft, 6890kg. There are several works, in the literature, that deal with aircraft modelling (see e.g. (19)). The main difference between the modelling of a fixed wing aircraft to another lies in the aerodynamics coefficients variation and the type of the employed propulsion. Hereafter, we recall the major modelling aspects of a fixed wing Aircraft. As shown in Fig.1, the roll\u2013pitch\u2013yaw convention are adopted using the Euler angles (\u03c6, \u03b8, \u03c8). Using Newton\u2013Euler convention, the dynamical equations of the aircraft are given by (1). We assume that the engines are positioned so that the thrust acts parallel to the aircraft body X -axis, with FT = l0\u03b4th and its point 2019 IFAC ACA August 27-30, 2019. Cranfield, UK 305 Copyright \u00a9 2019. Th Authors. Publi hed by Elsevier Ltd. All rights reserved. A. Hamissi et al. / IFAC PapersOnLine 52-12 (2019) 304\u2013309 305 ensure finite time convergence to the origin and grantee better accuracy (7)", + " In Section 3, the combining of HHOSM and NDI approaches for the autopilot design is addressed and HOSM differentiators are developed. The simulation environment and the results of implementation of this new autopilot are presented in Section 4. Finally, the conclusions are drawn in Section 5. In what follows, a brief description of the main features of the aircraft is introduced as well as its dynamic model. Then, the control problem is formulated. The chosen aircraft is a British Aerospace (Jetstream3102), which is a fixed wing, twin turboprop aircraft, illustrated in Fig.1. This type of aircraft has as control inputs the throttle setting command (\u03b4th), and the deflection angles of the three control surfaces: elevator (\u03b4e), ailerons (\u03b4a), and rudder (\u03b4r) (see Fig.1). We consider this aircraft due to the fact that all its aerodynamic coefficients are available in the literature (17),(18). The wing surface area S, the wingspan b, the mean aerodynamic chord c\u0304, and the mass m of the aircraft are considered as constant parameters given respectively as 280ft2, 46ft, 6.5ft, 6890kg. There are several works, in the literature, that deal with aircraft modelling (see e.g. (19)). The main difference between the modelling of a fixed wing aircraft to another lies in the aerodynamics coefficients variation and the type of the employed propulsion. Hereafter, we recall the major modelling aspects of a fixed wing Aircraft. As shown in Fig.1, the roll\u2013pitch\u2013yaw convention are adopted using the Euler angles (\u03c6, \u03b8, \u03c8). Using Newton\u2013Euler convention, the dynamical equations of the aircraft are given by (1). We assume that the engines are positioned so that the thrust acts parallel to the aircraft body X -axis, with FT = l0\u03b4th and its point 2019 IFAC ACA August 27-30, 2019. Cranfield, UK 306 A. Hamissi et al. / IFAC PapersOnLine 52-12 (2019) 304\u2013309 lies in the body-axes XZ-plane. The offset from the center of gravity is given by ZTP in the body-axes Z-direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001054_elan.201900109-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001054_elan.201900109-Figure2-1.png", + "caption": "Fig. 2. Schematic diagram of (a) the wall-jet cell (32 mm\u00d735 mm\u00d715 mm) with the clamp removed and (b) the flow path, which were not drawn to scale. Photographic images of (c) the flow channel, and (d), (e) the assembled flow cell. Alignment pins were used for the simple and precise construction of the flow cell. The substrate with the electrodes and the channel were fixed with a stainless-steel clamp.", + "texts": [ + " The PDMS channel was reversibly bonded onto the electrode substrate through manual alignment aided by a microscope. Lead wires were fixed at the contact pads of the electrodes with conductive adhesive. Figure 1 shows the design of the electrode system and a PDMS channel, as well as a photographic image of the resulting flow cell. www.electroanalysis.wiley-vch.de \u00a9 2019 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim Electroanalysis 2019, 31, 1\u20139 2 These are not the final page numbers! \ufffd\ufffd A flow cell with a PEEK block containing a channel was also constructed (Figure 2). An electrode-patterned substrate was fabricated on a PS substrate in the same manner as described above. The PEEK block was machined with a numerical control machining center in the workshop of Hokkaido University. The depth of the channel was 10 \u03bcm. Parallel pins were inserted into the PEEK block to align with the electrode substrate having the corresponding pin holes; the PEEK block and the substrate were fixed with an O-ring and SUS304 stainless steel clamps as shown in Figure 2. The inner diameters of the inlet and outlet were 300 and 1000 \u03bcm, respectively. All electrochemical experiments were performed with an ALS 1206A handheld electrochemical analyzer (BAS, Japan) at room temperature. Wall-jet cells were tested in flow-injection analysis (FIA) mode. The experimental setup included an MP 710i HPLC micropump (GL Science, Japan), a switching valve with a 2-mL inject loop, and the wall-jet cell in a Faraday cage (CS-2, BAS, Japan), as shown in Figure S1. For catechol, the carrier buffer was composed of 1 :2 : 7 methanol/0", + " We determined the optimum geometric configuration of the DEC-type ring array electrode to be that shown in Figure 7d (DEC2). Moreover, we also designed a wall-jet channel to evaluate the performance of the optimized electrode with a Ag/AgCl reference electrode. It was difficult to load a standard Ag/AgCl reference electrode into the PDMS channel employed in the above-mentioned studies. The channel was redesigned referring to a commercially available flow cell provided by BAS Inc., Japan and fabricated with PEEK resin and tightly clamped to the substrate with an O-ring, as shown in Figure 2. Figure S7 displays the fluid velocity profile in this flow cell as determined by numerical simulation, which demonstrates that radial flow forms in this flow cell. The dead volume in the flow cell was 0.7 pL, while the channel flow cell that we previously reported had a dead volume of ~20 nL [31]. The use of a Ag/AgCl reference electrode shortened the time for the baseline to stabilize to 2 min, which is to be compared with the PDMS-flow cell with a pseudo gold reference electrode that required over 1 h for the baseline to stabilize" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002960_physreve.101.062703-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002960_physreve.101.062703-Figure1-1.png", + "caption": "FIG. 1. Schematic illustrations of the LC cell with director profiles (a) before and (b) after the application of vertical electric field (E) for determining an unknown azimuthal anchoring energy (W) at homeotropic alignment layer.", + "texts": [ + " Here, several calculations should be carried out for many anchoring strengths to find the real anchoring strength that gives the desired midplane orientation (measured midplane orientation). We set the polar anchoring strengths by 5 \u00d7 10\u22125 N/m. In fact, the midplane direction remained unchanged at polar anchoring strengths, UT and UB above 10\u22125 N/m for a fixed azimuthal anchoring strength. III. EXPERIMENTAL METHOD The cell was assembled with two indium thin oxide (ITO)glass substrates that were coated with homeotropic and planar alignment layers. The alignment layers were rubbed perpendicular to each other [red arrows in Fig. 1(a)] to impose a hybrid-twist configuration of LCs as depicted in Fig. 1(a); 062703-2 planar anchoring at one substrate and homeotropic anchoring at the other substrate with 90\u25e6 twist of n across the cell. Under E, however, the LC configuration is transformed into a typical twist deformation due to \u03b5 < 0 of the used LC, MBBA is p-methoxybenzylidene-p-n-butylaniline (MBBA). Figure 1 shows the configuration of LC cells used in this work to determine the unknown W at homeotropic alignment surfaces. The LC cells were comprised of two bounding surfaces imposing a hybrid-twist LC (HTLC) configuration cell as shown in Fig. 1(a); planar anchoring at bottom substrate and homeotropic anchoring at top substrate with 90\u25e6 twist of n across the cell. Specifically, ITO-glass substrates were spin coated with a homeotropic alignment layer SE-1211 (Nissan Chemical) and a planar alignment layer of polyimide SE-3140 for HTLC-1 cells or polyvinyl alcohol (PVA) for HTLC-2 cells. The substrates were rubbed to set a preferred azimuthal orientation of LCs (i.e., easy axis) and assembled in a crossed fashion, Fig. 1(a). The thickness (d) of LC cells was set by glass spacers with a diameter of 8.5 \u03bcm. The rubbing conditions were as follows: the rotation speed of rubbing roll was 800 rpm, the rubbing depth was 3 mm, and the speed of rubbing stage was 5 cm/s. Subsequently, the experimental cells were filled with a nematic LC, MBBA which has \u03b5 < 0 (\u03b5\u2016 = 4.72 and \u03b5\u22a5 = 5.25), k11 = 6.66 pN, k22 = 4.2 pN, k33 = 8.61 pN, \u03b11 = \u221218.1, \u03b12 = \u2212110.4, \u03b13 = \u22121.1, \u03b14 = 82.6, \u03b15 = 77.9, \u03b16 = \u221233.6, and \u03c3||/\u03c3\u22a5 = 1.5 [36,40]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001296_aim.2019.8868485-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001296_aim.2019.8868485-Figure6-1.png", + "caption": "Figure 6. Single section bending diagram", + "texts": [ + " Through establishing a homogeneous transformation matrix of link parameters and joint variables, the relation between the terminal frame and the initial frame is deduced. The homogeneous transformation matrix is used to represent the relationship, so as to establish the motion equation of the continuum robot. In order to determine the relative relation between bending sections, frame {i-1} is attached to the beginning of the i-th section. The beginning of each section is attached to the ending of prior section, as shown in Fig.6. Because the axis of AEIR use rigid structure as a support, the axial is incompressible. Assuming the length of each part of AEIR is L, and each driving rope hole is distributed on the circumference of the radius r. Defining the angle between the curved plane of the continuum robot and the x-o-z plane is \u03b1, and the central angle corresponding to the single curved part is \u03b8. On the premise that the joint angle of each bending element is known, the length of each driving rope in the corresponding bending unit is as follows: 1 cosi i il L r = \u2212 (3) 2 cos( / 2)i i il L r = \u2212 + (4) 3 cos( )i i il L r = \u2212 + (5) 4 cos( 3 / 2)i i il L r = \u2212 + (6) Where lij represents the length of the j rope inside the pipe of segment i bending unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003601_1077546320954958-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003601_1077546320954958-Figure5-1.png", + "caption": "Figure 5. Map of accelerating test rig.", + "texts": [ + " The comparison of the lifetime parameter T with CT is shown in Figure 4. After feature extraction, GG fuzzy clustering is performed on the degradation division so that different degradation condition is able to be clustered with the fuzzy matrix calculated by the GG fuzzy clustering method. Finally, the clustering effect is evaluated by the three indicators expounded above. The whole lifetime dataset used in this section is from the IMS Center at the University of Cincinnati (Antoni and Borghesani, 2019). The test rig is shown in Figure 5. The testing bearing is Rexnord ZA-2115 double-row roller bearing, and the number of rollers is 16. The roller group pitch diameter is 75.501 mm, the roller diameter is 8.4074 mm, and the contact angle is 15.17\u00b0. One lifetime dataset is introduced for instance analysis which contains 984 groups of samples, and the final fault occurs at the outer ring. A time domain waveform is shown in Figure 6 in which the sampling interval has been filtered. It seems that the amplitude begins to increase at about the 700th group sample, and a quantitative feature needs to be extracted to describe the degradation process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000965_icma.2019.8816549-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000965_icma.2019.8816549-Figure2-1.png", + "caption": "Fig. 2 Schematic diagram of the optimal heading of the USV As shown in Fig. 2, the point cP is defined as the virtual center of the circular motion of the USV, and cr is the virtual radius. In the northeast coordinate system, the position of the catenary point is ( ),c c cP n e= , and the position of the USV is", + "texts": [], + "surrounding_texts": [ + "Since the area keeping of the USV can be regarded as a special positioning mode, which only requires the movement area of the USV no matter how to move, which direction it moves or even whether it moves in this area, we just make sure the USV is within the specified area. That is, the distance er between the USV and the regional center must satisfy Rer < , where R is the radius of the region. From the perspective of reducing energy consumption and increasing working time, this paper proposes a method for environmental optimal discontinuous control." + ] + }, + { + "image_filename": "designv11_80_0001380_012037-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001380_012037-Figure3-1.png", + "caption": "Figure 3. 3D models of (a) the SR mechatronic module, and (b) the robot.", + "texts": [ + " The markers are attached to the modules and combined into separate groups by the Vicon Tracker software. Each group corresponds to the head or the tail module (see figure 1). It is necessary to use at least three markers to determine the position of a group. A 3D model of the SR \u201cZmeelok-3M\u201d was created via SOLIDWORKS for virtual experimental study. The model was parameterized according to the design of the SR body elements and their inertial characteristics. Other parameters of the model match the specification (see table 1). Figure 3 shows an SR mechatronic module and the robot with 10 such modules as links. A computer model also created for virtual experiment includes: mathematical model of the SR dynamics, motor imitation model, mathematical model of the control system. A computer model was developed using MATLAB and Simulink and MSC ADAMS. The 3D model was imported from SOLIDWORKS Motion into MSC ADAMS as an assembly of solid bodies. Then it was parameterized and supplemented with a testing area model. Simulation of the dynamics of a mechanical system with unilateral contact was implemented in the MSC ADAMS" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure51.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure51.5-1.png", + "caption": "Fig. 51.5 Screen shot of microhole geometry measurement a diameter, b cylindricity", + "texts": [ + " P cos2hi P sinhi cos hi P coshiP sinhi cos h P sin2hi P sinhiP coshi P sinhi P 1 2 64 3 75 xo yo Ro 2 64 3 75 \u00bc P ri cos hiP ri sin hiP ri 2 64 3 75 \u00f051:6\u00de P cos2hi P sinhi cos hi P zi cos2 hi P zi sin hi cos hi P coshiP sinhi cos h P sin2hi P zi sin hi cos h P zi sin2 hi P sinhiP zi cos2 hi P zi sin hi cos hi P z2i cos 2 hi P z2i sin hi cos hi P zi sin hiP zi sin hi cos hi P zi sin2 hi P z2i sin hi cos hi P z2i sin hi P zi sin hiP coshi P sinhi P zi cos hi P zi sin hi P 1 2 6666664 3 7777775 xo yo lo mo Ro 2 6666664 3 7777775 \u00bc P ri cos hiP ri sin hiP rizi cos hiP rizi sin hiP ri 2 6666664 3 7777775 \u00f051:7\u00de 51 Characterization of Geometrical Features of Ultra-Short Pulse \u2026 609 The microhole\u2019s dimensions are measured at top, middle, and bottom sections by fitting three-point circle at the respective sections in the GOM software. The diameters of microhole are obtained as 240, 220, and 220 \u00b5m at top, middle, and bottom sections, respectively, with height of 0.056, 0.681, and 1.306 mm, respectively, as shown in Fig. 51.5. If the practice of quantifying taper based on entry (top) and exit (bottom) diameters is followed, there is a general taper in the microhole with the size narrowing with the depth. Similarly, the cylindricity error of microhole is measured as 11.2 lm by using Gaussian fit method, which is available in GOM Inspect software. The microhole geometrical xyz coordinate data given Table 51.1 is transformed to polar coordinates with respect to a reference circle. The reference circle is established by using a circle fitted through three selected points from given data", + "2, along with circularity errors and diameters. The circularity error is obtained as 3.182, 5.670, and 7.699 \u00b5m at top, middle, and bottom sections, respectively. From the above results, it observed that the quality of laser-drilled microhole changes with respect to depth. There is a good match between the diameter values obtained. The cylindricity of microhole is evaluated by considering the height (z-coordinates). The xyz coordinates are extracted from geometric model for top, middle, and bottom sections as shown in Fig. 51.5. Table 51.3 gives value parameters obtained in the cylindricity evaluation. The cylindricity error of microhole is evaluated as 610 K. Kiran Kumar et al. 51 Characterization of Geometrical Features of Ultra-Short Pulse \u2026 611 Gaussian fit method used in GOM Inspect, it is difficult to give reasons for the difference in the cylindricity errors. Interestingly, the diameter values agree with each other well. In the above table, subscripts R and o refer to reference cylinder and limacon cylinder, respectively, while x, y represent the intersection of axis with top plane; l, m are the slope of axis; r is the radius of the cylinder" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000095_s0025654418050047-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000095_s0025654418050047-Figure1-1.png", + "caption": "Fig. 1.", + "texts": [ + " It is shown that self-oscillating regimes arise in the supercritical region if the magnitude of the external torque exceeds a certain threshold value, while they are asymptotically stable for the whole range of their existence. 1. MECHANICAL MODEL AND EQUATIONS OF MOTION Consider a statically unbalanced rotor in the form of a hard disk fixed in the middle of a weightless f lexible shaft rotating on vertical hinged supports O1 and O2 under the action of constant external torque M. Using the Jeffcott model, we will consider the motion of the disk attachment point C in the horizontal S60 plane passing through the center of mass G of the disk. We introduce fixed OXY and rotating O\u03be\u03b7 coordinate systems, as shown in Fig. 1. By virtue of the assumptions made, the rotor has three degrees of freedom. For the generalized coordinates, we choose the absolute coordinates X, Y of point C and the angle of proper rotation of the disk \u03b8. The expressions for the kinetic and potential energy of the rotor have the form (1.1) where m is the mass of the rotor, IG is the moment of inertia of the disk about the axis passing through the center of mass G perpendicular to the plane of the disk, k is the coefficient of elasticity of the shaft, and s = |CG| is the value of static eccentricity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001137_kem.822.504-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001137_kem.822.504-Figure1-1.png", + "caption": "Fig. 1. Hybrid configuration on the left side and laser configuration on the right side", + "texts": [ + " Consequently, it has also been decided to assess the impact of pulsed and continuous laser radiation on the wall structure.The welding equipment is manufactured by EWM AG. The EWM Taurus 551 power source offers a welding current between 5 and 550 A and a welding voltage between 14.3 and 41.5 V for direct current GMAW application. The inert gas used was always 100% argon. For the manufacturing process of the samples further investigated in this work, two different configurations were used. These differ according to their type in hybrid process and laser process and in their geometry (Fig 1). In hybrid configuration the feeding angle is 90 \u00b0 and the laser-wire angle is 10 \u00b0. In pure laser mode configuration the feeding angle is 40 \u00b0 and the laser-wire angle is 65 \u00b0. The focus point of the laser was set to 20 mm and 30 mm above the substrate or the deposited layer for the tests in both configurations. Due to the two different configurations of the laser-GMA-hybrid test stand and the possibility to operate the laser in a modulated pulse mode as well as in CW mode, there are four different possibilities to generate additive manufactured walls" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003476_icsresa49121.2019.9182312-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003476_icsresa49121.2019.9182312-Figure1-1.png", + "caption": "Fig. 1. Representation of the five-phase permanent magnet synchronous machine with inter turns short circuit faults.", + "texts": [ + " In applications that don\u2019t require very high precision, the CMCM method is advantageous in terms of its simplicity of implementation based mainly on the fundamental laws of physics. This type of modeling offers a good compromise in terms of accuracy and calculation time. This method has already proved itself in the modelling of induction and synchronous machines dedicated to the detection of stator inter turns short-circuit fault[6 -8]. B. Modelling of the penta-phased synchronous machine by the CMC approach We consider a five-phase PMSM seat of five inter turns short-circuits in the stator as shown in Fig. 1. The five parameters (\u03b1a, \u03b1b, \u03b1c, \u03b1d, \u03b1e) define the percentages of shorted turns in relation with five respective faults. \u2022 Electric equations: After default the vector of the tensions given in equation (1) becomes: (1) Ts s s s s s s abcde a b c d e fv v v v v v v = Similarly, the currents vector (2) becomes: (2)s s s s s s s abcde a b c d e f T I I I I I I I= The new resistance matrix is expressed as follows: (1 )R 0 0 0 0 0 0 (1 )R 0 0 0 0 0 0 (1 )R 0 0 0 0 0 0 (1 )R 0 0 0 0 0 0 (1 )R 0 (3) 0 0 0 0 0 a a b b c c d d e e f R R \u03c3 \u03c3 \u03c3 \u03c3 \u03c3 \u2212 \u2212 \u2212 = \u2212 \u2212 The value of the resistance Rf is given by: " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure5-1.png", + "caption": "Figure 5. Generation of the shaft mesh", + "texts": [], + "surrounding_texts": [ + "with\ufeff{K}\ufeffis\ufeffrigidity\ufeffmatrix.\nMETHodS NUMERICAL ANALySES\nModel Presentation The\ufeffnumerical\ufeffsimulation\ufeffis\ufeffcarried\ufeffout\ufefffor\ufeffa\ufeffplain\ufeffbearing\ufeffof\ufeff100\ufeffmm\ufeffdiameter\ufeffand\ufefflength\ufeffof\ufeff80\ufeffmm,\ufeff the\ufeffmain\ufeffcharacteristics\ufeffof\ufeffwhich\ufeffare\ufeffshown\ufeffin\ufeffTable\ufeff1.\ufeffThe\ufeffbearing\ufeffis\ufeffembedded\ufeffin\ufeffa\ufeffsteel\ufeffring.\ufeffThe\ufeff pad\ufeffconsists\ufeffof\ufefftwo\ufefflayers,\ufeffthe\ufefflargest\ufefflayer\ufeffis\ufeffsteel\ufeffand\ufeff38\ufeffmm\ufeffthick,\ufeffthe\ufeffinner\ufefflayer\ufeffis\ufeffa\ufefftin-based\ufeff coating\ufeff(89%)\ufeffwith\ufeffa\ufeffthickness\ufeffof\ufeff2\ufeffmm.\ufeffThis\ufeffpad\ufeffis\ufefffed\ufeffby\ufeffa\ufefffeed\ufeffgroove\ufeffwith\ufeffa\ufefflength\ufeffof\ufeff70\ufeffmm\ufeff using\ufeffthree\ufefffeed\ufeffports,\ufeffdiameter\ufeff14\ufeffmm,\ufeff(Figure\ufeff3).\nTextures Parameters Surface\ufefftexturing\ufeffis\ufeffa\ufefftechnique\ufeffused\ufeffto\ufeffimprove\ufeffthe\ufeffload\ufeffcapacity\ufeffof\ufeffvarious\ufefftribological\ufeffconjunctions,\ufeff as\ufeffwell\ufeffas\ufeffto\ufeffreduce\ufefffrictional\ufefflosses.\ufeffThe\ufefftexture\ufeffis\ufeffspherical\ufeffshape,\ufeff(Figure\ufeff4)\ufeffwith\ufeffa\ufeffdiameter\ufeffrx\ufeff", + "=\ufeff3\ufeffmm\ufeffand\ufeffthe\ufeffdepth\ufeffof\ufeffry\ufeff=\ufeff0.5\ufeffmm,\ufeffthe\ufeffaxial\ufeffdistance\ufeffbetween\ufeffthe\ufefftextures\ufeffd\ufeff=\ufeff10\ufeffmm\ufeffand\ufefftheir\ufeff angular\ufeffoffsets\ufeff\u03b1\ufeff=\ufeff10\u00b0.\nMeshing The\ufefffinite\ufeffelement\ufeffnumerical\ufeffsimulation\ufeffis\ufeffused\ufeffto\ufeffcalculate\ufeffthe\ufeffdisplacement\ufeffof\ufeffthe\ufeffinner\ufeffface\ufeffof\ufeffthe\ufeff plain\ufeffbearing.\ufeffThe\ufeffsolid\ufeffis\ufeffdecomposed\ufeffinto\ufeffa\ufeffnumber\ufeffof\ufeff4-node\ufeffor\ufeff8-node\ufefftetrahedral\ufefffinite\ufeffelements\ufeff so\ufeffthat\ufeffthese\ufeffelements\ufeffare\ufeffas\ufeffaccurate\ufeffas\ufeffpossible\ufeffin\ufeffthe\ufeffgeometry.\nThe\ufeffshaft\ufeffis\ufeffdiscretized\ufeffinto\ufeffhexahedral\ufeffelements\ufeffwith\ufeff8\ufeffnodes,\ufeff(Figure\ufeff5),\ufeff15\ufeffnodes\ufeffin\ufeffthe\ufeffaxial\ufeff direction,\ufeff54\ufeffpoints\ufeffin\ufeffcircumferential\ufeffand\ufeff22\ufeffpoints\ufeffin\ufeffthe\ufeffradial\ufeffdirection.", + "The\ufeff pad\ufeff is\ufeff decomposed\ufeff into\ufeff 4\ufeff nodes\ufeff tetrahedral\ufeff elements\ufeff due\ufeff to\ufeff the\ufeff existence\ufeff of\ufeff a\ufeff groove\ufeffwith\ufefflines\ufeffand\ufeffarches\ufeffas\ufeffwell\ufeffas\ufefforifices\ufeffin\ufeffthe\ufeffcircle\ufeffform.\ufeffThese\ufeffshapes\ufeffrequire\ufefftetra\ufeff elements\ufeffto\ufeffachieve\ufeffthe\ufeffmost\ufeffaccurate\ufeffgeometry\ufeffoverlap,\ufeffFigure\ufeff6.\ufeffThe\ufeffbad\ufeffis\ufeffdecomposed\ufeffat\ufeff 12\ufeffpoints\ufeffin\ufeffthe\ufeffaxial\ufeffdirection,\ufeffthe\ufeffangular\ufeffamplitude\ufeffhas\ufeff70\ufeffpoints\ufeffand\ufeff4\ufeffnodes\ufeffdepending\ufeff on\ufeff the\ufeff thickness.\ufeff The\ufeff feed\ufeff groove\ufeff is\ufeff discretized\ufeff at\ufeff 13\ufeff points\ufeff in\ufeff the\ufeff radial\ufeff direction\ufeff and\ufeff 5\ufeffnodes\ufeffalong\ufeff its\ufeffwidth.\ufeffThis\ufeff finite\ufeffelement\ufeffmesh\ufeffof\ufeff the\ufeffplain\ufeffbearing\ufeffconsists\ufeffof\ufeff19830\ufeff elements\ufeffand\ufeff47784\ufeffnodes.\nBoundary Conditions To\ufeffbetter\ufeffresolve\ufeffthe\ufeffequations\ufeffpresented\ufeffearlier\ufeffin\ufeffthis\ufeffarticle,\ufeffit\ufeffis\ufeffnecessary\ufeffto\ufeffapply\ufeffappropriate\ufeff boundary\ufeffconditions.\ufeffThe\ufeffboundary\ufeffconditions\ufeffapplied\ufeffto\ufeffthe\ufeffshaft\ufeffare:\n\u2022\ufeff The\ufeffnodes\ufeffof\ufeffthe\ufeffface\ufeffwhere\ufeffthe\ufeffshaft\ufeffis\ufeffcoupled\ufeffwith\ufeffa\ufeffrotating\ufeffmanifold\ufeffare\ufeffblocked\ufeffalong\ufeffthe\ufeff x-axis,\ufeffFigure\ufeff7.\nThe\ufeffboundary\ufeffconditions\ufeffapplied\ufeffto\ufeffthe\ufeffpad\ufeffare:\n\u2022\ufeff The\ufeffpad\ufeffis\ufeffplaced\ufeffin\ufeffa\ufeffsupport\ufeffring\ufeffwhose\ufefflower\ufeffpart\ufeffis\ufeffspherical,\ufeffa\ufeffframe\ufeffone\ufeffbears\ufeffon\ufeffthe\ufeffface\ufeff of\ufeffthe\ufeffspherical\ufeffhydrostatic\ufeffplain\ufeffbearing\ufeffand\ufeffwhich\ufeffis\ufeffembedded\ufeffin\ufeffthe\ufeffbase,\ufeffFigure\ufeff8; \u2022\ufeff The\ufeffpad\ufeffis\ufefflocked\ufeffon\ufeff60\u00b0\ufeffof\ufeffthe\ufefflower\ufeffpart:\ufeffBlocking\ufeffof\ufeffthe\ufeffnodes\ufeffaccording\ufeffto\ufeffx,\ufeffy\ufeffand\ufeffz; \u2022\ufeff The\ufeffpad\ufeffis\ufeffmounted\ufeffin\ufeffa\ufeffring\ufeffso\ufeffit\ufeffis\ufeffcylindrical\ufeffsupport\ufeffwith:\n\ufeff\u25e6 Fixed\ufeffradial\ufeffnodes; \ufeff\u25e6 Free\ufeffAxial\ufeffNodes; \ufeff\u25e6 Fixed\ufefftangential\ufeffnodes.\nInsertion of Pressures We\ufeffcan\ufeffapply\ufeffglobal\ufeffloads,\ufeffstructural,\ufeffas\ufeffwell\ufeffas\ufeffimposed\ufeffdisplacements\ufeffaccording\ufeffto\ufeffthe\ufeffcases\ufeffstudied.\ufeff In\ufeffthe\ufeffcase\ufeffstudied,\ufeffpressures\ufeffare\ufeffapplied\ufeffalong\ufeffthe\ufeffcircumferential\ufeffaxis\ufeffas\ufeffwell\ufeffas\ufeffalong\ufeffthe\ufeffaxial\ufeffaxis\ufeff of\ufeffthe\ufeffshaft\ufeffand\ufeffthe\ufeffplain\ufeffbearing\ufeff(Figure\ufeff9)." + ] + }, + { + "image_filename": "designv11_80_0003579_978-981-15-8131-1_26-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003579_978-981-15-8131-1_26-Figure5-1.png", + "caption": "Fig. 5 Printing of nonplanar surfaces of test pieces with an infill angle of 45\u00b0. The image shows a 35\u00b0 part and the Zmorph nozzle depositing nonplanar filament lines on the inclination side upward vs. downward (red arrows). aDuring the upward filling, adjacent filament is slightly below currently deposited line, therefore resulting in minor scratching. b In a downward filling process, neighboring filament lines are slightly above the new line, which results in heavy scratching of them (red marks)", + "texts": [ + " In 3-axis nonplanar printing, it is not possible to adjust the extruder axis to be perpendicular to the printed surface. Therefore, the flat area at the nozzle tip plays an important role in the resulting surface quality. Measurements of the Zmorph\u2019s nozzle revealed that the diameter of the flat area is around 0.8 mm, which is twice the diameter of the filament extruding hole. This additional area is scratching the extruded filament in a downward printing motion. As the infill angle is 45\u00b0 in the printed examples, two of the sides are always printed slightly upward (Fig. 5a), and the other two are printed slightly downward (Fig. 5b). Please note that upward and downward refer to the direction in which the deposited layers are put next to each other. This leads to scratching of previously deposited filament when printing downward. Consequently, this results in a bad surface finish, which is clearly observable in the roughness of test pieces with an inclination of 30\u00b0. As the corner, where the printing of the last layer starts is defined as the front right, the left, and backside are the ones that are always printed downward" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003972_icma49215.2020.9233579-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003972_icma49215.2020.9233579-Figure3-1.png", + "caption": "Fig. 3 Mechanism for extension and returning of the wheel.", + "texts": [ + " The total weight of the robot body is 4kgf and the radius of the pinion gear is 1.3 cm. The minimum required torque is estimated to be 5.2kgf-cm, and the servo motor is selected to meet the requirement. The specification of the servo motor (ROBOT SERVO RS301CRF3 FUTABA) is shown in Table 1. 1) Mechanical Design of the Wheel We design the robotic wheel by using a CAD software. The design and appearance of the entire body is shown in Fig. 2. The foundation is designed with a rail structure that allows the spokes to slide stably as shown in Fig. 3. The picture of the assembled robotic wheel based on the design drawing is shown in Fig. 1. The robotic parts are printed by a 3D printer (AGILISTA-3200, KEYENCE) and also fabricated by a modeling plotter (NC-5SK, Mimaki). The materials of the parts are acrylate resin and ABS. 2) Rack & Pinion Mechanism A pinion gear connected with a servo motor is combined with a rack gear with a spoke as shown in Fig. 4. The extension can be achieved by the rack & pinion mechanism up to the half length of a radius" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000569_978-3-030-20751-9_33-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000569_978-3-030-20751-9_33-Figure3-1.png", + "caption": "Fig. 3: Architecture of the teleoperation system to command CDPRs", + "texts": [ + " This is because the pitch and roll of the trackball is directly reflecting the displacement of the planar motion of the CDPRs by integrating both sides of (6). The second scaling factor \u03bc2 dominates the perception of the motion of the CDPRs. In order to get the noticeable velocity perception on the CDPRs, \u03bc2 should be sufficiently large. By resisting the rotating trackball, the operator can decrease the speed of the trackball and consequently also the CDPR. 4 System Integration of a Dual Interface and CDPRs This section illustrates the system integration of the dual interface and CDPRs. The corresponding system architecture is depicted in Fig. 3. One challenge for the formulation of the dual interface is how tomerge two inputs of both the P2V joystick and the P2P trackball. In this system, the velocity is the common domain to blend two individual inputs to result in the reference CDPR trajectory through the blending policy (7). During teleoperation, the force of the joystick and the torque of trackball applied by the operator are measured at every time instant. Through the admittance controller, two different velocity inputs according to the above control laws in Sec" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002696_s40430-020-02368-5-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002696_s40430-020-02368-5-Figure4-1.png", + "caption": "Fig. 4 Principle of Mp", + "texts": [ + " Further investigation of the influence of key design parameters on stress and natural frequencies of gearbox is proposed for avoiding the resonance frequencies and reducing stress concentration. Figure\u00a03 shows the overall mechanical model of gearbox housing with external excitation caused by coal cutting. The regions of stress concentration and stress state of them are obtained by this model. The equations for gearbox housing are: where Mp is external torque acting on gearbox housing which is caused by meshing force between planet gears and ring gear, as shown in Fig.\u00a04. This torque was not considered by the previous papers [17\u201320, 23]. The principle of Mp is dynamic meshing force ( Fm ) which can be decomposed into radical ( Fy ) and circumferential directions ( Fx ) will be generated with the process of meshing between planet gears and ring gear. Circumferential direction force of gear meshing force is obtained, and Mp can be obtained.Mp , which equals Fx \u00d7 R , is then generated by Fx because of the bolt connections (shown in Fig.\u00a04) between ring gear and housing. It is noticed that other variables are listed in Table\u00a01. Section B\u2013B (Fig.\u00a01) which is located at the connection between motor shell and middle part of gearbox housing was identified as the unique region of stress concentration in the literature [23]. However, more fracture regions in section A\u2013A (Fig.\u00a01) which is located at the connection between middle part of gearbox housing and shell of planetary gear system were detected [2, 3] in engineering practice (Fig.\u00a05), and there are some field images about it in these literatures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001670_icems.2019.8922498-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001670_icems.2019.8922498-Figure2-1.png", + "caption": "Fig. 2. The Modal Test of Stator Core", + "texts": [ + " The elastic modulus of the loose assembly is much lower than that of the silicon steel sheet, so the effect of the winding in the stator\u2019s modal frequency is only on the quality. In the analysis, the winding is usually simplified to an additional mass loaded into the core slot. Taking the stator core of three frame series of Y-63, Y-100 and Y-160 as the experimental sample, the modal test model was established by LMS.Test.lab, and the unconstrained free mode of the stator core was tested by hammering method. state. The main dimensional parameters of the stator core test piece are shown in Table 1. The modal test condition is shown in Figure 2. The first 4th elastic mode frequency of these stator cores is shown in Table 2. The mode of each type of iron core is consistent, and the order of the mode shape of the Y-160 core mode is shown in Fig. 3. If the iron core is an isotropic material, the properties of the core material are the same as those of the 50WW470 silicon steel sheet (which elastic modulus is 2x10^11Pa, the Poisson's ratio is 0.28, and the density is 7850kg\u2022m^3), and the motor cores of three series of Y-63\u3001Y-100\u3001Y-160 are established" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003080_eucap48036.2020.9135366-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003080_eucap48036.2020.9135366-Figure1-1.png", + "caption": "Fig. 1. Structure of the proposed HWIR unit cell.", + "texts": [ + " An original and physical insightful circuit model was proposed, providing valuable explanations of the cell behavior for normal incidence. In the present contribution, we propose a single polarized HWIR (Half Waveguide Integrated Resonator) structure, additionally providing good angular stability. The manuscript is organized as follows. Section II describes the geometry and working principle of the HWIR. Section III introduces dielectric-loading strategy to improve oblique incidence stability. Finally, conclusions are drawn in Section IV. II. INITIAL DESIGN As it can be seen in Fig. 1, the proposed HWIR cell is composed of a section of rectangular waveguide of width Dx, height Dy and length t. Such a waveguide is loaded with a single metallic wire along z-axis, whose both ends are folded along y-axis. This resonator is thus sensitive to vertical polarization (along y). The geometry of the proposed vertically-polarized unit cell is inspired from the 3D-FSS in [6-8], and it has been conceived following the WIR perspective proposed in [9]. The operation principle of such a HWIR is based on the fact that impinging electromagnetic waves can be transmitted through the structure using two different paths: either through the waveguide section itself, i ither through the electric currents they induce on the wire resonator. The balanced control of these two alternative Authorized licensed use limited to: University of Glasgow. Downloaded on July 12,2020 at 12:06:05 UTC from IEEE Xplore. Restrictions apply. transmission mechanisms is key to the success in the synthesis of a given bandpass function. The cell periodicity, namely Dx and Dy (as shown in Fig. 1), determines the appearance of grating lobes (GL) due to the excitation of higher-order Floquet harmonics [REF]. In particular, Dy determines the GL excitation as the incidence angle increases in the Oyz plane. More precisely, the onset frequency of the first higher-order harmonic is given by: fonset = c0 /[p(1+sin )] (1) where c0 is the speed of light in free space, p is the unit-cell periodicity (p=Dy in this case), and is the incident angle. The original WIR cell is twice higher than the proposed HWIR cell since it is made of the symmetrical arrangement of two such cells, as depicts the Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002405_j.promfg.2019.07.037-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002405_j.promfg.2019.07.037-Figure3-1.png", + "caption": "Fig. 3. Optimized geometries for each minimum thickness studied on the bio-inspired 3D model.", + "texts": [ + " The smaller the minimum thickness (MT) constraint, the more precise it is going to be the geometrical optimization, as the software will be able to remove material up to that MT value. The MT constraint value should be as high as possible, in order to keep the run time as low as possible but being careful of maintaining a good enough resolution on the solutions. The value of 5mm has been used for the regular 3D model and bio-inspired 3D model for a more in deep comparison, as 5 mm presented the best compromise between time (Fig. 5) and structural significance of the final solution. (Fig. 3) The time required to perform all the geometrical optimizations has been obtained from the software timer. Three optimizations have been performed for each model, and then the average time and the standard deviation have been calculated. Errors smaller to 1% on the standard deviation have been neglected. Jaime Orellana et al. / Procedia Manufacturing 41 (2019) 121\u2013128 125 Author name / Procedia Manufacturing 00 (2020) 000\u2013000 5 A tetrahedral mesh has been automatically created by Inspire software" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003983_icma49215.2020.9233584-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003983_icma49215.2020.9233584-Figure7-1.png", + "caption": "Fig. 7 The flexible body model of tooth root crack failure", + "texts": [ + " When tooth root cracks of gears occur, the meshing force of the gears is the same as which in healthy planetary gearbox, so the simulation results cannot reflect the fault information. Thus, in the simulation of tooth root crack failure, the flexible deformation of fault gears should be considered. ABAQUS software is used to construct tooth root crack fault models of sun gear, planet gears and inner ring gear respectively. The crack propagation path is set as parabola [11]. The crack depth is 50% of maximum crack depth. After determining the element type and material properties, the mesh of crack area is divided finely, as shown in Fig.7 (a). The flexible body models of sun gear, planet gear and inner ring gear are shown in Fig. 7 (b), (c) and (d) respectively. Flexible body model of fault sun gear is used to replace the rigid sun gear in rigid-flexible coupling model of healthy planetary gearbox, so as to construct model of sun gear fault. Similarly, planet gear crack fault model and inner ring gear crack fault model can be constructed. 6 6.5 7 7.5 t(s) 20 10 0 -10 M ic ro st ra in 1965 Authorized licensed use limited to: Carleton University. Downloaded on June 15,2021 at 07:54:21 UTC from IEEE Xplore. Restrictions apply" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002057_s00006-019-1039-z-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002057_s00006-019-1039-z-Figure7-1.png", + "caption": "Figure 7. a Statics daughter hypervolume map (FV (p,f) [e1, e2, e3 ] |) of 3-RRR Manipulator. b Statics space (|fE | ) of the 3-RRR Manipulator", + "texts": [ + " The following kinematic pair parameters where chosen for this case study, where Hi is the initial vector of the kinematic pair from joint i to i+1, (in meters), and si is the initial unit vector along the joint axis: H1 = {0.5, 0, 0} H2 = {0.5, 0, 0} H3 = {0, 0.5, 0} H4 = {0, 0.5, 0} H5 = {0,\u22120.5, 0} H6 = {0,\u22120.5, 0} s1 = {0 , 0, 1} s2 = {0, 0, 1} s3 = {0 , 0, 1} s4 = {0 , 0, 1} s5 = {0 , 0, 1} s6 = {0 , 0, 1} Kinematic inputs were chosen to be (rad/s and rad/s2) \u03b8\u03071 = 1.5 \u03b8\u03072 = 1.5 \u03b8\u03073 = 1.5 \u03b8\u03081 = 1 \u03b8\u03082 = 1.5 \u03b8\u03083 = 1 \u03c4max = 25Nm Comparing Fig. 7a, b against each other and similarly with Fig. 8a, b we note that there is a remarkable similarity in structure between the statics and space functions and their daughter hypervolume map counterparts. The qualitative determination of the functions being geometrically similar is clear. Using Eq. (35.9c), we found that the approximation strength index of the forward statics and kinematics cases are R(forward statics \u2212 3RRR) = 0.910 and R(forward kinematics\u22123RRR) = 0.906 which reflects that the functions are scaled representations of each other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000194_012042-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000194_012042-Figure1-1.png", + "caption": "Figure 1. Tested pneumatic spring: (a) \u2013 scheme of experimental pneumatic spring; (b) \u2013 general view of experimental unit on the stand; 1 \u2013 sleeve type RCC; 2 \u2013 hydropulsator with support plate 3; 4 \u2013 bored piston; 5 \u2013 sprung mass (SM); 6 \u2013 adjusting channels; 7 \u2013 upper cover; 8 \u2013 filler connector; 9 \u2013 bore of RCC 1; 10 \u2013 bore of piston 4; 11 \u2013 body of air damper; 12 \u2013 clamp; 13 \u2013 bolts; 14 \u2013 sealing; 15 \u2013 throttle; 16 \u2013 radial ports; 17 \u2013 elastic non-return valve; 18 \u2013 buffer of maximum compression; 19 \u2013 manometers; 20, 28 \u2013 valves; 21 \u2013 compressor; 22 \u2013 bypass valve; 23 \u2013 check valve; 24 \u2013 filter; 25 \u2013 intake; 26 \u2013 force sensor; 27 \u2013 position sensor; 29 \u2013 connecting pipe with receiver 30; 31 \u2013 displacement and speed sensor; 32 \u2013 hose; 33 \u2013 guides; \u041d \u2013 height of PS in static position; \u03c2 \u2013 displacement of the hydropulsator plate; CU \u2013 compressor unit", + "texts": [ + " In order to assess heating rates the Chair of \u201cAutomated units\u201d of the Volgograd State Technical University performed bench tests of the pneumatic spring with and without air damper at a different operational volume. The purpose of these tests was to define the zones of maximum heating for different PS surfaces, as well as the peak temperature and the time for reaching a stabilized temperature mode. Tests have been performed on a dynamic stand with servo-hydraulic actuator co-developed with Biss ITW (India) (figure 1) [15]. The test PS 1 is installed on a bench between the support plate 3 of hydropulsator 2 and adjusting channels 6 of the sprung mass 5. The test PS has a sleeve type rubbercord casing (RCC) 1 VL 260-340, a bored piston 4 and an upper cover 7 with a filler connector 8. The bore of RCC 1 is connected with a bore 10 of piston 4 through the air damper with body 11 axially installed on the top end of piston 4 and fixed using a clamp 12 and bolts 13 with a sealing 14. Inside the body 11 there is a throttle 15 interconnecting bores 9 and 10, two rows of radial ports 16 and elastic non-return valve 17 shutting down these ports during the rebound and connecting bore 9 with bore 10 of the piston 4 only during the compression" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003818_d0sm01339k-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003818_d0sm01339k-Figure2-1.png", + "caption": "Fig. 2 Helix angle - Trimers align along each other with a preferred curvature R\u03ba = sin2 \u03a8 and twist R\u03c4 = sin\u03a8cos\u03a8. Due to their embedding in the curved surface of the cylinder this leads to an average bond orientation with the helix angle \u03a8.", + "texts": [ + " The trimer-trimer interactions consist of two pair potentials: the LJ potentials VLJ(ri j), and the elastic bending and torsion potential Vel(ri j), Vpair = \u2211 i, j ( VLJ(ri j)+Vel(ri j) ) (1) VLJ(ri j) =V AA LJ (rAA i j )+V BB LJ (rBB i j )+V AB LJ (rAB i j ), (2) with the distance vector r\u03b1\u03b2 i j = r\u03b1 i \u2212 r\u03b2 j between the center of the corresponding particles of the trimers i and j. The distance between two trimers is measured as the distance between the attractive beads, ri j = rBB i j . The location of the trimers is restricted to the surface of a cylinder of radius R (see Fig. 2). We assume that the self-assembled filament behaves like a elastic twisted ribbon on length scales larger than the trimer units, with an elastic deformation energy for bending and torsion22,27. Thus, two trimers on the curved membrane surface, which are closely bound together, acquire an elastic deformation energy Vel = [ B\u03c3 (\u03ba\u2212\u03ba0) 2 +T \u03c3 (\u03c4\u2212 \u03c40) 2 ] S(|~ri j|), (3) when the distance d = |ri j| between their center B-particles is smaller than the elastic interaction range a = 1.25 \u03c3(< rB c ). Here, \u03ba and \u03c4 are the curvature and twist of the assembly, respectively, broken down to the bond level, and \u03ba0 and \u03c40 are the corresponding intrinsic, preferred values", + " I, measured in LJ units, except for the length scale, which measured in units of the elastic binding distance a. The trimer density \u03c1 and the self-diffusion coefficient D define a time scale t0 = 1/(\u03c1D). This time scale corresponds to the average time it takes a trimer to diffuse the average nearest-neighbor distance. All simulations are run for about 13\u00d7106 t0. A snapshot of the current configuration is stored every 12\u2032000 t0. Particles are considered to belong to the same cluster if ri j \u2264 1.5 \u03c3 . If not stated otherwise, the configuration of the last frame of the simulation is analyzed, see Fig. 2 as an example. Particle positions on the cylindrical hyperplane are mapped to x and y coordinates, with periodic boundary conditions in both directions. The mapping allows the reconstruction of curvature \u03ba and twist \u03c4 of the bonds, as discussed below in detail. The main axis of the cylinder is oriented along the x-axis, whereas the angular coordinate \u03d5 \u2208 [0,2\u03c0] is unwrapped along the y-axis, i.e. y = \u03d5R \u2208 [0,2\u03c0R]. We employ elongated cylinders with a constant length-to-radius ratio L/R = 25. There is an optimal number N\u2217 of trimers, which allows for exactly one perfect fitting helix, with N\u2217 = L/(\u03c3 cos\u03a8) = 25R/(\u03c3 cos\u03a8), for a given helix angle \u03a8 (Fig. 2). This corresponds to an optimal packing fraction of p\u2217f (\u03a8) = N\u2217 AT /A = At/(2\u03c0R\u03c3 cos\u03a8), where At = 2.89 \u03c32 is the area covered by a single trimer, and A is the area of the simulation box. We simulate at two different packing fractions p f = NAT /A, which are p f = 0.15 and p f = 0.21. For a cylinder radius of R = 62.5 \u03c3 and the smaller (larger) packing fraction the simulated system contains a total of N \u2248 32 (45)\u00d7 103 trimers for a box area A = 392,500 a2. For smaller radii the trimer number is adapted accordingly", + "],1\u201311 | 3 S of tM at te r A cc ep te d M an us cr ip t Pu bl is he d on 1 2 O ct ob er 2 02 0. D ow nl oa de d by U ni ve rs ity o f W es te rn O nt ar io o n 10 /1 3/ 20 20 1 1: 15 :4 2 A M . fold22. This implies on a cylindrical manifold that the orientation of a bond with respect to the cylinder axis corresponds to a certain curvature and twist. The formation of clusters of several trimers leads to helix formation, where each bond orientation angle corresponds to the local helix angle \u03a8 (see Fig. 2). On a cylindrical manifold, a bond with angle \u03a8 directly implies a local curvature R\u03ba = sin2 \u03a8 and twist R\u03c4 = sin\u03a8cos\u03a8 = (1/2)sin(2\u03a8), as derived by Andrews and Arkin22. This can be seen best by considering the bond as a part of a helix, wrapped around the cylinder of radius R, with an helix angle \u03a8, where tan\u03a8 = dy/dx = R/p \u2208 [\u2212\u03c0/2,\u03c0/2], and p is the helix pitch. We can then reformulate the elastic potential (3) in terms of the helix angle \u03a8, Vel = B\u03c3 ( sin2 \u03a8 R \u2212\u03ba0 )2 +T \u03c3 ( sin\u03a8cos\u03a8 R \u2212 \u03c40 )2 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001251_embc.2019.8857814-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001251_embc.2019.8857814-Figure1-1.png", + "caption": "Figure 1: Mechanical design", + "texts": [ + " Stewart platforms have now been found a wide variety of applications, such as, flight simulators, parallel robot manipulators, machine tools, 6 degrees of freedom coordinate measuring devices, entertainment and health equipment. Stewart platforms inherent advantages over the conventional serial mechanism are simpler structure, higher stiffness, better accuracy, and heavier loading ability. It mainly consists of a lower base platform, an upper mobile platform (\u00d81100 mm) and six identical stretchable legs (see Figure 1). Each end of the legs is mounted to the platforms with two cardan joints (see Figure 1 Details A and B). Each leg is actuated by a linear actuator (Rexroth Bosh group, Germany). Features of the motors we adopted are summarized in Table 1. The platform is mounted on an actuated rotating support for enhancing the limited yaw range because of the Stewart configuration (1S-series OMRON Corporation, Japan). Above the mobile platform, immediately after the cardan joints, we inserted on the passive limbs amplified load cells (TCE-amp AEP Transducer, Italy) whose nominal load has been specially lowered to 100 kg, see Figure 1 Detail C. In order to facilitate the use with patients the plate of the platform was mounted at the floor level taking advantage from the availability of a room under the Movement Analysis Laboratory (MARlab). The platform is first elevated only when a translation on horizontal plane is necessary, avoiding impacts with the adjacent floor. The robot can be controlled in position and force, or it is possible to generate trajectories and/or control the force of the platform as a function of external feedback signals, such as the platform torque computed by the outputs of the load cells" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000821_ccaa.2018.8777578-Figure17-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000821_ccaa.2018.8777578-Figure17-1.png", + "caption": "Fig. 17: Top view of the arena 3.", + "texts": [ + " In this arena the quad was tested for the horizontal zigzag motion along the row of pillars and to test the response when the control instructions changes instantaneously. This arena has three pillars in a row with each pillar of a circumference of 119.65 cm and 365.76 cm apart. Fig. 16: Practical implementation of the arena 2. Table II shows the number of trials performed and time measured for the completion of horizontal zigzag motion. The best result obtained in this arena was in 44 sec. The Quad has good response to instantaneous changes in controls. Arena 3: Fig. 17 shows the top view of the arena 3 and Fig. 18 shows the practical implementation of the arena 3.The arena consists of a circle placed a height of 548.64cm from the ground. This arena was made to test the altitude stability. TABLE III: OBSERVATION OF THE ARENA 3. Trial Time(S) 1 15 2 17 3 12 4 15 5 14 Table III shows the number of trials performed and time taken to complete the task. This test is performed to find quad\u2019s altitude stability, as the quad need to be in the level of the circle to pass through it" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001467_012017-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001467_012017-Figure1-1.png", + "caption": "Figure 1. Axis at Ordinate 1 at 2.08s, 6.66s, and 10 s", + "texts": [ + " Variations in this study based on the axis of the water wheel axis at 1 meter, 1.25 meters, and 1.5 meters. From each variation will be analysed for 10 seconds. But in the results will only be shown 4 time is 0.4 seconds, 2.08 seconds, 6.66 seconds, and 10 seconds. Where in this simulation used the initial velocity of water on the river is 1.2 \ud835\udc5a/\ud835\udc60 and move faster based on time. The simulation will be shown in the picture below. ICETsAS 2018 Journal of Physics: Conference Series 1376 (2019) 012017 IOP Publishing doi:10.1088/1742-6596/1376/1/012017 From the simulation FIGURE 1, result at 1-meter axis ordinate, the maximum speed of water flow based on the simulation result is 34.6 \ud835\udc5a/\ud835\udc60. And the most stable flow is shown at 6.6 seconds. Based on the simulation FIGURE 2, result at the time of the ordinate of the shaft 1.25 meter, the maximum speed of water flow based on the simulation result is 28.4 \ud835\udc5a/\ud835\udc60. And the most stable flow shown at 6.6 seconds is also the same when the ordinate of the shaft is 1 meter. Based on the simulation FIGURE 3, result at the time of the ordinate of the shaft 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000530_978-3-030-12082-5_55-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000530_978-3-030-12082-5_55-Figure7-1.png", + "caption": "Fig. 7 Distribution of contact pressure on contact surface of platform, Pa", + "texts": [], + "surrounding_texts": [ + "A number of computational experiments were carried out, the conditions and results of which are summarized in Table 2. The temperature field of the macromodel was determined, and then, to compare experiment results, the thermal contact conductance of the gyro unit-platform contact region averaged over the nominal area was calculated by the relation \u03b1c q\u0304 T2 \u2212 T1 where q\u0304 is the average heat flux density over the area of finite elements of the lower nominal surface, T1, T2 are the average temperatures over the area of finite elements of the nominal lower and upper contact surfaces, respectively. Experiments were conducted to evaluate the influence of various factors, while the most appropriate model should be considered as the experiment No. 4 model. The TCC parameter was set as a constant value equal to 157,000 W/(m2 K) or as the above-tabulated dependence on pressure TCC(p). The value 157,000 W/(m2 K) is obtained from the graph in Fig. 5 for the average pressure from the bolt clamp force of 7.3 MPa, calculated by dividing the sum of the clamp forces of each bolt (2000 N) by the nominal contact area. To evaluate the effect of thermal expansion, experiments were carried out for two types of contact behavior, Standard and No separation. For the Standard contact type, the contact heat transfer occurred strictly in the real contact area (Fig. 6), for which the TCC parameter was set. Thus, this type of contact reflects the influence of change in shape from thermal expansion. For the No separation contact type, movement of the contact surfaces along the contact plane is allowed, but separation of the surfaces is not permitted and the real contact area is equal to the nominal one. Thus, in the case of No separation contact, the change in shape of the surfaces from thermal expansion is not reflected in temperature results since it does not affect the thermal contact conductance. In this case, the heat transfer occurs over the entire nominal contact area. The wide use of this type of contact in actual practice is due to the significantly lower computational complexity and, accordingly, solution time. The No separation contact type was set on the gyro unit-platform connection in experiments Nos. 1 and 2. The calculations were carried out with the assumption of small displacements, since the accounting for large displacements for experiment No. 6 resulted in a change in the averaged thermal contact conductance of 0.1%, which is considered insignificant. The first and second experiments set the heat transfer throughout the whole of the nominal contact area. In this case, setting the dependence of the thermal contact conductance TCC on the contact pressure p obtained in the micromodel led to a decrease of 4.3 times in the averaged thermal contact conductance of themacromodel. The models used in the computational experiments Nos. 3 and 4 take into account the effect of thermal expansion on the real contact area. Because of the change in shape of the cylindrical body of the gyro unit, tangency takes place in the form of a narrow ring along the outer edge of the nominal contact area.Also, areas near the bolts are in direct contact. The real contact area was 56% of the nominal area (Fig. 6). As is clear from a comparison of experiments Nos. 3 and 1, the averaged thermal contact conductance decreased bymore than 5 times just due to accounting for the real contact area at a constant TCC of 157,000 W/(m2 K). Under the same conditions and using the TCC(p) dependence (experiments Nos. 2 and 4), the averaged thermal contact conductance decreased noticeably less, by 56%, which can be considered a result of thermal expansion without the direct influence of contact pressure. Repetition of the result of 56% is a random coincidence in this case. Experiment No. 5 showed that the use of a friction coefficient 0.3 instead of 0.5 led to a slight increase in the thermal contact conductance (by 18%). Thus, the friction coefficient has a noticeable effect on the conductance of the actual contact. The thermal contact conductance is significantly affected by clamp force of the bolts. Experiment No. 6 showed that using the conditions of experiment No. 4 and decreasing the clamp force from 2000 to 100 N, the averaged thermal contact conductance decreased by more than 6 times. A similar effect in real structures can arise in the case of more complicated connections, for example, with clasps [35]. The distribution of contact pressures, surface temperatures, and heat fluxes for the contact platform surface in experiment No. 4 is shown in Figs. 7, 8, and 9. Figure 10 shows the distribution of temperatures throughout the entire model of the gyro unit-platform assembly under the same conditions." + ] + }, + { + "image_filename": "designv11_80_0001299_aim.2019.8868869-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001299_aim.2019.8868869-Figure2-1.png", + "caption": "Figure 2. Piezo-actuated jet dispensing valve profile", + "texts": [ + " Finally, a piezo-actuated jet dispensing valve based on the proposed FHDA is developed, and corresponding experimental platform and control system are built in our lab. Fig. 1 shows the straight circular bridge displacement amplification mechanism used in this paper. The internal piezoelectric actuator generates micro-displacement as input, and its output direction points to the outer displacement to produce amplification effect. The 3D profile of the piezo-actuated jet dispensing valve, which a FHDA is mounted to implement displacement amplification is shown in Fig. 2. Experimental Investigation of Piezo-actuated Jet Dispensing Via a Flexible Hinge Displacement Amplifier Proceedings of the 2019 IEEE/ASME International Conference on Advanced Intelligent Mechatronics Hong Kong, China, July 8-12, 2019 978-1-7281-2493-3/19/$31.00 \u00a92019 IEEE 1613 Equations (1) and (2) are the different reference formulas for the FHDA magnification ratio derived from this paper. Its derivation and parameter types can be referenced paper[14]. 1 2 2 2 1 2 3 ( ) _ 3 w l L f A Qi w f s t f + = + (1) 3 2 2 1 2 2 3 2 2 2 1 6 ( ) 3 _ (2 ) (6 3 ) h l L f R L w A Lin hf L R h h f R L w + + = + + + (2) A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000546_978-3-030-20131-9_278-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000546_978-3-030-20131-9_278-Figure2-1.png", + "caption": "Fig. 2. Waist mechanism with 3 DOFs: (a) initial position and (b) working position.", + "texts": [ + " Modern waist joint/structures contain reliable and low backlash mechanisms, such as epicyclic gear trains, harmonic planetary gears, harmonic drives and cycloidal reducers, enabling high positioning accuracy and repeatability of movements which is essential for motion control [19]. However, deficiency of these mechanisms is a very high price that limits their use. For this reason, we suggest waist mechanism with simple structure, based on the lever mechanisms enabling high carrying capacity and low backlash. Figure 2 shows the 3 DOFs waist mechanism that allows pitch, yaw and roll rotations. It consists of three interconnected segments: bottom, middle and upper disc \u2013 links 1,6,8 respectively. Link 1 is immovable and has the absolute coordinate system Oxyz positioned at point O. Vertical link 1b is fixed to link 1 at point O and it represents the first fork of cross 7. This cross is connected to link 1b by the cylindrical joints \u2013 points P and Q are centers of joints. The second fork is integrated in link 6 and is connected to cross 7 with the cylindrical joints \u2013 points R and S are centers of joints", + " The lever \u2013 link 10, is connected by rotational joints to links 9 and 8 in points M and N, respectively. The coordinate system O2x2y2z2 is fixed to link 8 in point O2 \u2013 in the initial position, the axes of the coordinate system O2x2y2z2 are parallel to the axes of the absolute coordinate system Oxyz. In addition, the axes of the coordinate systems O2x2 1y2 1z2 1 and O2x2 2y2 2z2 2 are parallel with the axes of the coordinate systems O1x1 1y1 1z1 1 and O1x1 2y1 2z1 2. Due to the transparency of the Fig 2b, the coordinate system O2x2 1y2 1z2 1 is not shown. Finally, after rotation around the z2 2 axis for the angle 8 6, the coordinate system O2x2 2y2 2z2 2 switches to the position O2x2 3y2 3z2 3. Following, based on known displacements of the input links 2, 4 and 9, by applying direct kinematics, rotation angles and angular velocities of the output links 6 and 8 are determined. Figure 3a shows the mechanism 1a,2,3,6 for the rotation of link 6 around the axis defined by points P and Q. The rotation angle of link 6 is determined according to: ( ) ( )( ) ( ) ( ) 2 22 1 0 2 1 0 2 6,1/2 2arctan OO VA s OO VA s A AB OV A AB OV \u03d5 \u2212 + \u00b1 \u2212 + \u2212 + + = \u2212 + (1) where: ( ) ( )( )222 2 1 0 2 1 12 BF AB OV OO VA s O F A O F \u2212 + \u2212 \u2212 + \u2212 = (2) The angular velocity of link 6 is determined according to: ( ) 2 6 1 6 6 3cos sin cot s O F \u03d5 \u03d5 \u03d5 \u03d5 = \u2212 (3) where: 1 6 3 cos arccos O F AB OV BF \u03d5\u03d5 \u2212 \u2212 = (4) Figure 3b shows the mechanism 1c,4,5,6 for the rotation of the link 6 around the axis defined by points R and S" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002057_s00006-019-1039-z-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002057_s00006-019-1039-z-Figure6-1.png", + "caption": "Figure 6. a The 3-RRR Manipulator; b the network representation of a 3-RRR Manipulator in terms of enumerated kinematic pairs", + "texts": [ + " This generates a single physical solution as expected. The parallel kinematics expression in accordance with [2,35] is given as (as shown in Eq. (29)) 1X \u03b8 kn+1 \u03b8\u03071 = 2X \u03b8 ln+1 \u03b8\u03072 = \u00b7 \u00b7 \u00b7 = nX \u03b8 pn+1 \u03b8\u0307n = TE (37) For serial kinematics, we have [2,35] TE = [\u2211n i=1 X\u03b8 i n+1, \u2211n i=1 (UGi siU \u22121 Gi ) ] \u03b8\u0307i (38) and the expression for serial statics becomes (adapted from [35]) \u03c4i = X\u03b8 in+1 \u00b7 fE + si \u00b7 nE (39) 3.1.1. The 3-RRR Manipulator Forward Statics and Kinematics Comparison. In this example, the 3-RRR manipulator\u2019s (Fig. 6) statics and kinematics daughter hypervolumes for the [e1, e2, e3 ] bases, FV (p,f) [e1, e2, e3 ] and FV (p,k) [e1, e2, e3 ] (from Eqs. (32) and (25) respectively), are compared against the forward static and kinematic end effector outputs of |fE | and |vE | respectively, where FV (p,k) = Min[F1V (s,k)\u03b31, F4V (s,k)\u03b34, F3V (s,k)\u03b37] (40a) F1V (s,k) = 1X \u03b8 1,10 \u2227 1X \u03b8 2,10\u03b32 \u2227 1X \u03b8 3,10\u03b33 (40b) F2V (s,k) = 2X \u03b8 4,10 \u2227 2X \u03b8 5,10\u03b35 \u2227 2X \u03b8 6,10\u03b36 (40c) F3V (s,k) = 3X \u03b8 7,10 \u2227 3X \u03b8 8,10\u03b35 \u2227 3X \u03b8 9,10\u03b39 (40d) and FV (p,f) = V {1, 2} \u03b41\u03b42 2 + V {1, 3} \u03b41\u03b43 2 + V {2, 3} \u03b42\u03b43 2 1 (40e) where X \u03b8 i n+1 is computed according to Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure83.13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure83.13-1.png", + "caption": "Fig. 83.13 Three mode shapes of the adapter", + "texts": [ + " shows that as load is increased, the displacement also increases in the negative z-direction, i.e. along the height of the adapter and also follows a linear relationship. The deformed shape of the adapter superimposing onto the undeformed configuration is shown in Fig. 83.12. The results obtained from static analysis have been saved and further utilized for extraction of mode shapes in modal analysis. In modal analysis, three modes were extracted and the results obtained in modal analysis are as follows (refer Fig. 83.13): The modal frequencies are listed in Table 83.4. Fig. 83.11 Displacement in z-direction at different loads Fig. 83.12 Deformed shape versus original shape The results from different models are studied and found that if there is increase in the cross section of the helical groove of carbon fibre, the buckling factors increase. Also, if two circumferential ribs are used, it contributes in increasing buckling factor. It can also be noted that, the bottommost layer in the ply sequence is the critical layer in the helical ply as it is experiencing the highest value of stress (in both the ribs circumferential and helical plies)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003225_03772063.2020.1779136-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003225_03772063.2020.1779136-Figure1-1.png", + "caption": "Figure 1: Introducing the coordinate axes of the fuselage", + "texts": [ + " This choice, however, was one of the most important challenges in doing this research. The process of this paper is as follows, after the introduction and the works related to nonlinear modeling of rotorcrafts as \u00a9 2020 IETE well as the extraction of the linear model, LPV control is introduced with its corresponding model in the introduced system and expression of equations in this field. Finally, the simulation of different scenarios in tracking the routes will be presented. The axes of coordinates of the fuselage would be introduced before introducing the nonlinear model. Figure 1 illustrates how these axes are introduced and the forces acting on it. The self-propelled rotorcraft image is used in [28]. It is introducing the unmanned rotorcraft control inputs that have four inputs. It can be stated that one of these four inputs is related to the tail rotor and three of them are related to the main rotor. The control inputs are considered as normalized between (\u22121, 1). Typically, general equations of motion can be stated that one of these four inputs is related to the tail rotor and three of them are related to the main rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001090_032034-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001090_032034-Figure3-1.png", + "caption": "Figure 3. The minimum and maximum allowable values of the indicators of the size of the total contact patch% H - along the height of the tooth,% L - along the length of the tooth.", + "texts": [ + " The satellite billet is installed by the auto operator into the expanding horizontal mandrel, the broach rotates around the vertical axis and receives a coordinated translational movement parallel to the line of the depressions of the gear wheel. Both machines are structurally similar, the difference lies in the performance of interchangeable sections of circular broach and copiers, defining the trajectory of their relative movement relative to the stretched workpiece. In the process of toothing, it is necessary to ensure that the profile and the relative position of the side surfaces of the teeth of the ring gear so that the contact patch is located at the calculated point of the gear train [1] (fig. 3). In the workshop conditions, under the supervision of the personnel, there is only one quality indicator of the satellite - the beating of the ring gear along the pitch circle, which is the only indicator for evaluating the technological accuracy of the drive-through automatic line, for measurements of which there is a special control device at the workplace. Evaluation of the size of the contact patch is carried out by visual comparison of the obtained contact patch with the allowed forms. Measurements and registration of its parameters is not performed", + " 2), as a result of which it is unsuitable for developing comprehensive recommendations for quality improvement [6] Therefore, before developing a method for controlling the quality of the manufacture of gear rims in the toothing process, a method was developed to assess the compatibility of the gear teeth of the differential satellite in terms of both dimensions and the relative position of the total contact patch. As a result, in addition to the standard indicators characterizing the size of the contact patch relative to the length and height of the tooth% H,% L (fig. 3), the indicators L1, L2, L3, L4 were added - the distances from the extreme points of the total contact patch to the borders of the active tooth surface (fig. 4, 5). Mechanical Science and Technology Update IOP Conf. Series: Journal of Physics: Conf. Series 1260 (2019) 032034 IOP Publishing doi:10.1088/1742-6596/1260/3/032034 To find the whole complex of operating factors affecting the magnitude of the total contact patch, taking into account the results of a positive solution of similar problems in various processes of forming machine parts [7\u201311], the main key quality indicators were found, the values of which should be measured when production experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure1-1.png", + "caption": "Figure 1. Modeling strategies [2].", + "texts": [ + " The first step necessary for such a study was quantification of the deviation of the kinematic output of a mechanism with clearance from that of an ideal mechanism. The appropriate statistical tool for*For correspondence this has been identified and used for analyzing all the simulations. Modeling of joints with clearance is very crucial to analyze the performance of mechanisms. With the help of these models, the designer can estimate the output and corrective actions can be taken to reduce the output error. The commonly used modeling strategies for mechanisms are shown in figure 1. a) The clearance is modeled by adding a virtual massless link that has a constant length equal to the radial clearance. b) Spring-damper can also be used to model the joint clearance. c) The journal and bearing are considered as two colliding bodies known as momentum exchange approach. The third approach is more realistic for a dry clearance joint. It allows the contact force models to be applied and it takes into account the dissipation of energy during the impact process. Radial clearances introduce two extra degrees of freedom in the mechanism" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003334_speedam48782.2020.9161955-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003334_speedam48782.2020.9161955-Figure1-1.png", + "caption": "Fig. 1: Reconfigurable multi-motor drive in which each motor can be connected or disconnected to a gear mechanism", + "texts": [ + "ndex Terms\u2014induction motor, Kalman filter, rotor flux observer, torque estimation, torque control, system identification I. INTRODUCTION Nowadays, many induction motor drives require a torque control, for example electric vehicles or multi-motor drives which are considered in the given application. Within a multimotor drive system (comp. Fig. 1), the mechanical power is not provided by one but multiple motors which are connected reconfigurable by a gear mechanism. Such a drive is normally operated in its entirety in speed-control mode whereas each motor operates in torque-control mode [1]. If there are high deviations between the actual torque values and their reference values, inefficiencies or even control problems of the complete drive might arise. In standard field-oriented control (FOC) schemes which are due to its simple structure and its good performance still very popular in modern industrial drives, the torque is controlled open-loop (comp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002609_dese.2019.00022-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002609_dese.2019.00022-Figure3-1.png", + "caption": "Fig. 3. Klann-based limb. a) free-body diagram. b) Klann limb with motor.", + "texts": [ + " The presented approach provides simulations that concern the underactuated mobility space of a hexapod. In section II, the proposed underactuated mechanical design and the Klann limb kinematics are presented. In section III presentes the deduction of the dynamic control walking model and section IV provides the conclusion. The present work proposes a locomotion mechanism consisting of a differential driving system (Figure 2a), an all-limb synchronous bidirectional steering yaw mechanism (Figure 2b) and the Klann-based limbs (Figure 3). The right and left sided are differential speeds providing instantaneous velocity vt and yileding instantaneous yaw speed, \u03c9t. One drive per lateral triplet of limbs, interconnected front/back sides by tracks from the central drives. The steering mechanical system of Figure 2b shows bilateral direction angle for all-limb synchronously and can turn in yaw \u2212\u03c0/4, ..., \u03c0/4. The Klann mechanism (Figure 3) is an underactuated planar multi-link system, which from rotary input motion, it produces as output a cycloid trajectory (Figure 4). The links proportions are defined to optimize linear motion of the contact point at every rotary half cycle of the crank. The contact point lifts during the other rotary half cycle, before returning to the staring position. Klann [9] presented his famous invention, the \u201dwalking device\u201d, which is usually deployed to emulate biological limbs motion, mostly antropods. For instance, [21] developed artificial active whiskers for underwater guidance of walking robots based on the Klann mechanism. The Klann mechanism has several functional advantages as numerous advanced displacement systems have, such as stepping over obstacles, climb stairs, walking over all terrains, while not requiring a computing system to be controlled. The Klann mechanism is comprised of seven rigid links (L1, ..., L7, Figure 3a), a stretcher frame, a link used as crank handle L1, two as seesaw that are joined to the stretcher frame (L2 and L5), and all of them are interconnected through pivot joints (A,B, ..., F ). The links proportions define the optimize the link-foot linearity of motion with half rotation of the crank handle (Figure 3b). The rest of the crank\u2019s rotation allows that the link-foot move up to a defined height before motion return toward the staring position, repeating again the motion loop. In this work, a Klann mechanism of specific proportions was designed for the purpose of limbs in a tripod-like walking. The kinematic model was deduced to numerically simulate and analyze the gait tracks. The link L1 in connection with L3 actuate as a rotating cam with L2 as a balancing bar. The oscillatory motion is transmitted to the coupled links L4, L5 and L6 to convert the rotary motion of L1 into linear motion in the base of the link effector L7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002215_s00170-019-04806-8-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002215_s00170-019-04806-8-Figure4-1.png", + "caption": "Fig. 4 Projections of velocities of planar potion on the fixed coordinate axes \u041e1x, \u041e1y", + "texts": [ + " Velocities of symmetric points of contact \u04121 and \u04122 are equal to point \u0412 velocity, since they are located on one perpendicular line to the symmetry plane, passing through point \u0412, hence, the values of the contact points\u2019 velocities shall be equal to V\u03b3 B1 \u00bc V\u03b3 B2 \u00bc V\u03b3 B \u00bc BL \u03b3 ; V\u03b3 K \u00bc KL \u03b3 : Projections of velocities V\u03b3 B1; V \u03b3 B2;V \u03b3 K of points \u04121, \u04122, and \u041a on movable axes \u041e2\u03b7 and \u041e2\u03b6 are equal to (see Fig. 3) V\u03b3 B1\u03b7 \u00bc V\u03b3 B2h \u00bc \u2212BL \u03b3 cos\u03b52 \u00bc \u22122a2 \u03b3 ; V\u03b3 B1\u03b6 \u00bc V\u03b3 B2\u03b6 \u00bc \u2212BL \u03b3 sin\u03b52 \u00bc \u22122S2 \u03b3 : V\u03b3 K\u03b6 \u00bc KL \u03b3 ;V\u03b3 K\u03b7 \u00bc 0: \u00f010\u00de Projections of these velocities V\u03b3 B1;V \u03b3 B2;V \u03b3 K to auxiliary axis \u041e1h are equal (see Fig.4) to V\u03b3 B1h \u00bc V\u03b3 B2h \u00bc \u2212BL \u03b3 cos\u03b51 \u00bc \u22122a1 \u03b3 ; V\u03b3 Kh \u00bc \u2212KL \u03b3 sin\u03b3: \u00f011\u00de These expressions include (see Fig.3) \u03b51\u2014angle between segment \u0412L and axis of hole, \u03b52\u2014angle between segment \u0412L and peg axis, cos\u03b51 \u00bc a1 OB; sin\u03b51 \u00bc S1 OB; sin\u03b52 \u00bc S2 OB; cos\u03b52 \u00bc a2 OB; B \u00bc OB \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a21 \u00fe S21 q \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a22 \u00fe S22 q ; BL = 2OB = 2B. Projections of velocities V\u03b3 B1; V\u03b3 B2; V \u03b3 K to fixed axes\u041e1\u0445, \u041e1\u0443, and \u041e1z (Fig. 4) are equal to V\u03b3 B1x \u00bc V\u03b3 B2x \u00bc \u2212V\u03b3 B1hcos\u03c8 \u00bc \u22122a1cos\u03c8\u03b3 ; V\u03b3 K\u0445 \u00bc \u2212KLsin\u03b3cos\u03c8\u03b3 ; V\u03b3 B1y \u00bc V\u03b3 B2y \u00bc \u2212V\u03b3 B1hsin\u03c8 \u00bc \u22122a1sin\u03c8\u03b3 ; V\u03b3 Ky \u00bc \u2212KLsin\u03b3sin\u03c8\u03b3 ; V\u03b3 B1z \u00bc V\u03b3 B2z \u00bc \u22122S1 \u03b3 : V\u03b3 Kz \u00bc KL \u03b3 cos\u03b3: \u00f012\u00de Rotation about hole axis occurs with angular velocity of \u03c8\u0307 \u00bc d\u03c8 dt . Velocities of contact points in this movement are located in the fixed plane \u041e1\u0445\u0443, directed (see Fig.5) at tangents to the hole aperture edge circumference and at all contact points are equal in magnitude V\u03c8 B1 \u00bc V\u03c8 B2 \u00bc V\u03c8 K \u00bc 0:5D\u03c8\u02d9 : Projections of these velocities to the fixed axes of coordinates \u041e1\u0445 and \u041e1\u0443 (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003517_1464419320955114-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003517_1464419320955114-Figure6-1.png", + "caption": "Figure 6. Elliptical contact area at (a) number of rolling elements 5, (b) number of Rolling elements 6.", + "texts": [ + " The lowest quantity of balls (Nb\u00bc 5) carries more load on the individual rolling element where the highest numbers of balls (Nb\u00bc 11) distribute the less loading value. Therefore, the majority bearing manufacturers optimize their curvature ratio in the range of 0.51 to 0.54. As per the hertzian contact theory, an elliptical contact area is occurred due to individual load on balls. The results are elaborated by considering an axial Load 6000N by varying the curvature ratio from 0.51 to 0.58. In all plots the maximum contact area generated due to minimum curvature ratio (fi\u00bc 0.51) and maximum contact area occurs due to maximum curvature ratio (fi\u00bc 0.58). Figure 6(a) indicates the contact area of five numbers of rolling elements. The smallest contact area is 1.26 mm2 and the maximum area occurred is 1.94 mm2. There is a 53% expansion in the contact area for the range selected for the minimum to maximum curvature ratio at constant load. When the smallest curvature ratio (fi\u00bc 0.51) is considered, this creates a close surface contact and a comparatively excess surface in touch with a rotating element. When six numbers of balls are chosen, the minimum contact area is 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002846_012072-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002846_012072-Figure8-1.png", + "caption": "Figure 8. Phenomenon on tire tread: (a) Before the main plate moves; (b)When the plate is moved at a distance x, the tire tread also deform as long as x.", + "texts": [ + "1088/1757-899X/788/1/012072 Based on figure 7, the length of contact patch can be calculated with equation (2). (2) where lt is a contact-patch length, rf is unloaded tire radius, y is the difference between unloaded and loaded tire radius (rl). The value of y is measured by LVDT 1. When the load W is applied to the wheel, the tire will deflect in vertical direction. Because of that, the wheel centre will move down at distance y from the original position. The relationship between slip and plate displacement is shown in figure 8. Figure 8(a) shows the wheel condition when the plate is not moved yet. Figure 8(b) is the condition when the plate is moving at distance x. The wheel is locked. It means that the wheel does not rotate. In this research, the tire characteristic modelling is done only from a slip which equals to zero until a slip at maximum friction force. Therefore, it can be assumed that the total deformation happened along contact patch lt and will be the same as the plate displacement x. The is a part of the tire tread affected by deformation (stretching). ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10.1088/1757-899X/788/1/012072 The effect of tire tread deformation toward distance travelled by the wheel is shown in figure 9. From figure 9(a) it can be seen that when the plate is moved, an angle appears. This angle is also shown in figure 8(b). Thus, if the wheel rotates with an angle , it will travel like if it rotates with an angle + . Therefore, the distance travelled by the wheel will be further compared to the wheel with no deformation . This effect also can be explained by figure (b). Deformation S makes the effective radius (re) longer than the free rolling radius (r0). As a result, with the same angular velocity, the braked wheel will travel further than the non-braked wheel. Slip can be defined as the percentage of distance difference travelled by a wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.10-1.png", + "caption": "Figure 1.10. Left: a view of the rotation of angle # around n; right: a visualization of the section corresponding to the Rodrigues rhombus", + "texts": [ + " Deduce from this that the matrix associated with the linear operator Rn,% is written as: Rn,% = $ % 1 0 0 0 1 0 0 0 1 & ' +(1\" cos&) $ % \"n2 y \" n2 z nxny nxnz nxny \"n2 x \" n2 z nynz nxnz nynz \"n2 x \" n2 y & ' +(sin&) $ % 0 \"nz ny nz 0 \"nx \"ny nx 0 & ' . 3) Conversely, we are given a rotation matrix Rn,% for which we wish to find the axis of rotation n and the angle of rotation &. Compute the trace tr (Rn,%) and Rn,% \" RT n,% and use it to obtain n and & in function of Rn,%. For a geometric illustration, see Figure 1.10. 4) Deduce a method to find a interpolation trajectory matrix R(t) between two rotation matrices Ra, Rb such that R(0) = Ra and R(1) = Rb. 5) Using a Maclaurin series development of sin& and cos&, show that: Rn,% = exp (& \u00b7Ad (n)) , which sometimes written as: Rn,% = exp (& \u00b7 n') . EXERCISE 1.6.\u2013 Quaternions Quaternions, discovered by W.R. Hamilton, are also used to represent a rotation in a three-dimensional space (see, e.g. [COR 11]). A quaternion q\u030a is an extension of the complex number. It corresponds to a scalar s plus a vector v of R3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003327_speedam48782.2020.9161896-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003327_speedam48782.2020.9161896-Figure9-1.png", + "caption": "Fig. 9. Contour diagram of the stress at 17000 rpm.", + "texts": [ + " Therefore, it is necessary to discuss the mechanical strength in the proposed motor. The yield point stress of rotor core used in the proposed motor is 275 MPa. When the safety factor is 1.5, the allowable maximum stress is about 183.3 MPa. Consequently, the maximum speed is the speed when the stress applied to the rotor reaches 183.3 MPa. Fig. 8 shows the stress for the speed in proposed motor. As is obvious from this figure, the stress in the rotor shows the maximum allowable one at the speed of 17000 r/min. Fig. 9 shows the contour diagram of the stress applied to the rotor at 17000 r/min. Fig. 10 shows the efficiency map of proposed model. Fig. 11 shows the efficiency map of model A. According to the results, the proposed model shows 70% or more under 10000 r/min, while the conventional model A shows 70% or more under 9000 r/min. Expansion of the high efficiency range is due to the high electrical resistivity of the bonded neodymium magnet. 659 Authorized licensed use limited to: McMaster University. Downloaded on October 17,2020 at 20:29:58 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001726_icems.2019.8921944-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001726_icems.2019.8921944-Figure1-1.png", + "caption": "Fig. 1. The structure of permanent magnet auxiliary SynRM", + "texts": [ + " While the magnetic circuit model analysis and finite element simulation are carried out, the test method of dynamic demagnetization of the motor is proposed innovatively, and the effectiveness of demagnetization evaluation is improved. II. MOTOR STRUCTURE AND DEMAGNETIZATION PROBLEM The PMASyn-RM has two forms: distributed winding and concentrated winding. The concentrated winding is widely used in small power motors due to its small end and small copper loss. This paper is mainly for the arrangement of arc-shaped magnetic steel. The permanent magnet auxiliary synchronous reluctance motor is analyzed, and its structure is shown in Fig. 1. The running vector diagram of the PMASyn-RM is shown in Fig. 2. Since the control adopts Id<0, The directaxis magnetic potential of the motor is the demagnetizing magnetic potential of the permanent magnet. At the same time, the ferrite permanent magnet has a low coercive force. When the direct-axis weak magnetic current of the motor is large, the permanent magnet is subjected to the demagnetizing magnetic potential, therefore, in combination with the operation of the motor, it is necessary to evaluate the anti-demagnetization ability of the motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000778_j.jsc.2018.02.004-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000778_j.jsc.2018.02.004-Figure1-1.png", + "caption": "Fig. 1. Grid-like frameworks in (R2, \u2016 \u00b7 \u2016\u221e), where one of the vertices is fixed at the origin: the framework in (a) has two degrees of freedom, as p1 and p2 can move vertically and horizontally, respectively, independent of each other; the framework in (b) has one degree of freedom, as p2 can still move horizontally; the framework in (c) is rigid. The colours of the edges are induced by their orientation relative to the unit ball in (R2, \u2016 \u00b7 \u2016\u221e).", + "texts": [ + " Published by Elsevier Ltd. This is an open access article under the CC BY license (http://creativecommons.org/licenses/by/4.0/). A bar-joint framework in the plane is referred to as grid-like if the bar-lengths are determined by a norm with a quadrilateral unit ball. The allowable motions of such a framework constrain vertices adjacent to any pinned vertex to move along the boundary of a quadrilateral which is centred at the pinned vertex and obtained from the unit ball by translation and dilation (see Fig. 1). This is an important context from the point of view of applications. For example, the problem of maintaining rigid formations of mobile autonomous agents is a well-known application of geometric rigidity theory and its associated \u201cpebble game\u201d algorithms (see Eren et al., 2004). However, the Euclidean metric may not always be the most natural choice for controlling a formation. For instance, it may not be possible to detect Euclidean distances between agents (e.g. due to obstacles in the terrain)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000542_978-3-030-20131-9_261-Figure20-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000542_978-3-030-20131-9_261-Figure20-1.png", + "caption": "Fig. 20. Beaded string process state diagram of the uplink line penetrating into the common beads.", + "texts": [ + " Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state. Based on the problem existed for the hand-weaving method, the automatic weaving method and mechanism device are studied and discussed. The weaving route was extracted by discussing the hand-weaving method of the beaded cool pad" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001549_s11223-019-00110-8-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001549_s11223-019-00110-8-Figure13-1.png", + "caption": "Fig. 13. Thermal strain cloud atlas of the wave generator.", + "texts": [ + " Because thermal elastic deformation occurs on the basis of the initial deformation, the two \u201csprings\u201d are superimposed on each other. Therefore, the integrated meshing stiffness becomes \u201cflexible,\u201d and its structural sketch is shown in Fig. 17. Integrated meshing stiffnesses kH of the wave generator and movable tooth and kK of ring gear and movable tooth can be calculated as follows: k k k H HT HN 1 1 1 , (19) k k k K KT KN 1 1 1 . (20) Fig. 14. Thermal strain cloud atlas of the movable tooth. Fig. 13 Fig. 14 Fig. 15 Fig. 16 Fig. 15. Thermal meshing stiffness of the wave generator and movable tooth. Fig. 16. Thermal meshing stiffness of the movable tooth and ring gear. The integrated meshing stiffness data of the meshing pairs are plotted in Figs. 18 and 19. The maximum value of kK of the ring gear and movable tooth is 3.75 10 7 N/m. Value of kK exhibits a sudden jump near 50 , mainly due to the change of kKN . After considering the thermal strain, a 31.4%-reduction of the meshing stiffness of the wave generator and movable tooth was observed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000609_978-3-030-12684-1_17-Figure17.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000609_978-3-030-12684-1_17-Figure17.1-1.png", + "caption": "Fig. 17.1 Solid ABS part dimensions", + "texts": [ + " Test specimens used for the series of experiments herein are built with ABSplus-P340 polymer using a uPrint SE Plus [30], or with 316L stainless steel using a Concept Laser M2, which is a direct metal laser sintering machine [31]. The parts are designed to be rectangular cantilevers with high aspect ratio cross-sections to obtain relatively large displacements normal to the largest face and provide a wide, flat surface convenient for DIC speckling. Three solid ABS parts are printed with the dimensions shown in Fig. 17.1. Each part is built in a different orientation, as shown in Fig. 17.2. That is, all three parts in this set have the same outer geometry, but have their build layers oriented in different Cartesian directions. A second set of three ABS parts is printed with internal lattices. Each lattice is a 2D cell pattern printed along the build orientation, as shown in Fig. 17.3. Thus the second set of parts corresponds to the same build orientations as in Fig. 17.2, with lattices oriented vertically with respect to the base plate. The lattices are only built in the cantilever section of each part\u2014the base is left solid. The 2D cell pattern consists of 1.5 mm square holes separated by 1 mm solid walls, and surrounded by a 1 mm thick external wall. Figure 17.4 shows the outer dimensions of the lattice ABS set, which are increased from that of the solid ABS parts to allow for the internal lattice structures. A third set of parts is built in solid steel. These parts are built with the same geometry as the solid ABS parts (Fig. 17.1), except they are 3 mm thick instead of 4 mm, and have a 5 mm fillet instead of 4 mm. The three steel parts are also built in the same orientations as the three solid ABS parts (Fig. 17.2). The parts were excited by the translational oscillations of a shaker table using the Shaker Control software from Bruela & Kjaer Sound and Vibration Measurement. The shaker table was oriented so that the large face of the part oscillated normal to the imaging plane. Parts are secured to the shaker table with 4 mm bolts, as seen in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000303_978-981-13-3627-0-Figure9.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000303_978-981-13-3627-0-Figure9.1-1.png", + "caption": "Fig. 9.1 Picking-up of thin chip from the substrate: a practical operation of thin-chip transferring with different vacuum picking-up heads, b schematic illustration of a large thin-chip bending under the vacuum adsorption through a single hole, c the chip is kept flat under the condition of the vacuum adsorption through distributed hole-matrix. It should be noted that the top surface of the vacuum ejector cap is flat to tightly hold the substrate, which can establish a flat contact between the substrate and the cap. \u00a9 2015 IEEE. Reprinted, with permission, from Ref. [10]", + "texts": [ + " D\u2019Couto GC, Babu SV (1994) Heat transfer and material removal in pulsed excimer-laserinduced ablation: Pulsewidth dependence. J Appl Phys 76(5):3052\u20133058 Chapter 9 Vacuum-Based Picking-up and Placing-on After partly being peeled off by the peeling-off processes discussed in previous chapters [1\u20133], the chip picking-up plays a critical role in flip-chip packaging to detach the chip intactly from the substrate. An illustration of the chip picking-up process through the vacuum adsorption is shown in Fig. 9.1. It is difficult to peel thin chip directly off from the substrate, which is restricted by a potential risk of the chip damage [4\u20137]. Whether the thin chip can be detached intactly from the remaining adhesive region or not is determined by the capability of the picking-up with the vacuum adsorption [8, 9]. With the increase of the ratio of the chip length to the chip thickness, the resulting incomplete detachment even leads to the chip failure. In order to determine a proper process window, it is necessary to explore effects of main process parameters on the chip picking-up process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001467_012017-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001467_012017-Figure3-1.png", + "caption": "Figure 3. Axis at Ordinate 1.5 at 2.08s, 6.66s, and 10 s", + "texts": [ + "1088/1742-6596/1376/1/012017 From the simulation FIGURE 1, result at 1-meter axis ordinate, the maximum speed of water flow based on the simulation result is 34.6 \ud835\udc5a/\ud835\udc60. And the most stable flow is shown at 6.6 seconds. Based on the simulation FIGURE 2, result at the time of the ordinate of the shaft 1.25 meter, the maximum speed of water flow based on the simulation result is 28.4 \ud835\udc5a/\ud835\udc60. And the most stable flow shown at 6.6 seconds is also the same when the ordinate of the shaft is 1 meter. Based on the simulation FIGURE 3, result at the time of the ordinate of the shaft 1.5 meter, the maximum speed of water flow based on the simulation result is 26.1 \ud835\udc5a/\ud835\udc60. And the most stable flow shown at 6.6 seconds is also the same at the time of the ordinate of the shaft is 1 meter and 1.25 meters. From the three variations can be seen, that the flow of fluid in the river, stable or not, based on the travel time of water (probably due to the same water wheel shape) is 6.6 seconds. As for the fluid velocity in the river is more influenced by the axis ordinate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002906_pen.25409-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002906_pen.25409-Figure13-1.png", + "caption": "FIGURE 13 Predicted parison shape as a function of extrusion time at Weissenberg number Wi = 40 and for a die exit cross-sectional aspect ratio \u03c7 = 62.5, using a viscoelastic PTT model: front view of the extruded sheet showing sinusoidal wrinkle deformation", + "texts": [ + " The magnitude of such stresses that cause wrinkling depends solely on the magnitude of the normal stress difference developed in the die. The minimum compressive stress induced in the extruded sheet soon after exit from the die is N1 = \u22120.065 MPa (see Figure 9D). As this compressive stress is below the critical compressive stress (\u03c3c = \u22120.0524 MPa) for the extruded sheet to wrinkle, that is, N1 < \u03c3c (note that we are comparing negative stresses), wrinkling occurs, in accordance with the energy method presented previously. Predicted extrudate wrinkles at various times during extrusion are shown in detail in Figure 13 (front view) and in Figure 14 (perspective view). By comparing these results with the previous ones (see Figure 7) obtained at same Weissenberg number (Wi = 40) but with a smaller die exit cross-sectional aspect ratio (\u03c7 = 12.5), we observe the absence of wrinkles for a smaller aspect ratio. All these findings are consistent with the energy method which shows that the critical compressive stress \u03c3c for sheet wrinkling scales like 1/\u03c72. In the particular case of extrusion of a Newtonian fluid from a rectangular die with a large aspect ratio \u03c7 = 62" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001881_s12206-019-1113-4-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001881_s12206-019-1113-4-Figure4-1.png", + "caption": "Fig. 4. Bending fatigue test machine.", + "texts": [ + " We performed gas carburizing as a case-carburizing method of the gears, as shown in Fig. 3. First, we conducted the carburizing process for 240 min under the carburizing temperature 930 \u00b0C and the carburizing carbon potential Cp = 1.25 %. Second, we conducted the diffusion process for 180 min under the diffusion temperature 930 \u00b0C and the diffusion carbon potential Ck = 0.85 %. Then, we cooled the gears at 830 \u00b0C for 30 min, oil-cooled them at 50 \u00b0C, and finally aircooled them. Note that similar case-carburizing processes have been reported [6, 27]. Fig. 4 schematically shows the bending fatigue testing machine used in our bending fatigue test [3, 6]. This machine can engage a test gear and a supporting gear at an arbitrary meshing position, and apply a cyclic load to the gear pair. This machine was composed of a fuel injection pump generating hydraulic pressures, a pump drive unit, a pressure controller, a torque generator, and a gear supporting unit. We mounted the test gear (drive gear) on the same shaft as the torque generator, and fixed both shaft ends of the supporting gear (driven gear)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001843_j.ifacol.2019.12.321-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001843_j.ifacol.2019.12.321-Figure3-1.png", + "caption": "Fig. 3. Forces on the fins due to the flow and flapping", + "texts": [ + " * B f (9) Here, 1 1 2 2 3 3 1 1 3 32 2 3 3 3 31 1 1 1 2 1 2 2 3 3 3 32 2 2 2 1 1 1 1 2 2 2 2 3 3 3 3 1 1 1 1 2 2 2 1 1 1 1 0 0 0 0 0 0 ) ) ) ) - - - * *- * * * * ( * ( *( * ( * * * * * * * ( * ( * * * * *) ) s c s c s c s c s cs c y c y sz c z s y c y s z s z cz s z c z s z c x c x c x c x s x c x s y s y y s y c B 2 3 3 3 3- * *c y s y c and T 1 1 2 2 3 3f D L D L D L where \u03b2 is the angle at which the flow is striking on the fins relative to the vertical. The vertical axis for the caudal and pectoral fin is the y and z-axis respectively, as shown in Fig. 3. The fin oscillation angle \u03b3 is also expressed relative to the vertical axis. \u03b1 is the angle of attack which gives the flow direction with reference to the fin given by \u03b1 = \u03b2 \u2013 \u03b3. The subscript 1 represents the parameters of the caudal fin, and 2 and 3 are that of the pectoral fins respectively. Thus, the transformation of lift and drag forces into force/moment vector is expressed as The summation of the rigid body term CRB(\u03bd) and the added mass term CA(\u03bd) gives the Coriolis and centripetal matrix C(\u03bd) which is parameterized as skew-symmetrical", + " * B f (9) Here, 1 1 2 2 3 3 1 1 3 32 2 3 3 3 31 1 1 1 2 1 2 2 3 3 3 32 2 2 2 1 1 1 1 2 2 2 2 3 3 3 3 1 1 1 1 2 2 2 1 1 1 1 0 0 0 0 0 0 ) ) ) ) - - - * *- * * * * ( * ( *( * ( * * * * * * * ( * ( * * * * *) ) s c s c s c s c s cs c y c y sz c z s y c y s z s z cz s z c z s z c x c x c x c x s x c x s y s y y s y c B 2 3 3 3 3- * *c y s y c and T 1 1 2 2 3 3f D L D L D L where \u03b2 is the angle at which the flow is striking on the fins relative to the vertical. The vertical axis for the caudal and pectoral fin is the y and z-axis respectively, as shown in Fig. 3. The fin oscillation angle \u03b3 is also expressed relative to the vertical axis. \u03b1 is the angle of attack which gives the flow direction with reference to the fin given by \u03b1 = \u03b2 \u2013 \u03b3. The subscript 1 represents the parameters of the caudal fin, and 2 and 3 are that of the pectoral fins respectively. Thus, the transformation of lift and drag forces into force/moment vector is expressed as *Sin *cos ; *Sin *cos x i i i i z y i i i i F D L F or F L D Fig. 3. Forces on the fins due to the flow and flapping The propulsive forces of the fins are calculated along the xbaxis for the caudal fin and along yb,-axis and zb-axis for the port and starboard pectoral fins respectively. The total force components generated for propulsion are shown in equation (10). Similarly, the moments for propulsion about the same body axes are given in equation (11). The frame orientation for the caudal fin and pectoral fins are shown in Fig. 2 and Fig. 3 respectively. The force along zaxis and moment about the z-axis is equal to zero for the caudal fin and force along y-axis and moment about the yaxis is equal to zero for both the pectoral fins. *Sin *cos *Sin *cos 3 x j j j j j 1 3 z y j j j j j 1 F D L F or F L D (10) *Sin *cos ; *Sin *cos *Sin *cos ; *Sin *cos 3 x j j j j j j 1 3 y j j j j j j j j j j j 1 3 z j j j j j j 1 M y L D M x L D z D L M y D L (11) A nonlinear controller is designed to track the desired trajectory, to hold the desired position and to stabilize the AUV with a minimal error which requires good robustness to handle the uncertainties despite the occurrence of external disturbances and internal noise" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000724_978-981-13-3305-7_190-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000724_978-981-13-3305-7_190-Figure5-1.png", + "caption": "Fig. 5. The transition stages demonstration", + "texts": [ + " Besides, more effective and powerful propulsion system is needed to obtain the larger tangential acceleration of the flight path to reach the minimum flight speed of fixed-wing mode in a relatively short time. In this paper, the discontinuous transition strategy is used because of its easy realizability and simple flight parameter adjustment procedure, and the continuous transition strategy will be determined in the future. There are mainly three stages in the discontinuous transition flight process shown in Fig. 5, that is, the initial wing deflection stage, the acceleration stage and the fixedwing preparation stage. A more specific description is listed in the Table 2. The reasons why there are three stages are listed below: Firstly, when the parameters of the propulsion system are fixed, it is impossible for the aircraft to deflect its wing from the vertical state to the horizontal state directly, because there exists the minimum airspeed requirement when the aircraft enter the fixed-wing flight mode. Thus, before entering the fixed-wing mode, there should exist enough time for airplane to obtain the minimum airspeed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000629_ceit.2018.8751827-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000629_ceit.2018.8751827-Figure3-1.png", + "caption": "Fig. 3. Rotor loops", + "texts": [ + " Local defects on the outer, inner raceways and ball cause vibrations whose frequencies f0 , fi and fb , are calculated as follows [5]: Outer raceway: fc=f0= nn 2 fr 1- Db Dc cos \u03b2 (1) Inner raceway: fc=fi= nn 2 fr 1+ Db Dc cos \u03b2 (2) Ball: fc=fb= Dc Db fr 1- Db 2 Dc 2 cos2 \u03b2 (3) Characteristic mechanical vibration frequencies described by (1) to (3) can be seen in the stator current spectrum of the motor [6]. This leads to the following frequencies in stator current: fbf=fs\u00b1k fc (4) I. MATHEMATICAL MODEL OF DSIM The stator of the motor considered in this study has six phases divided into two sets of symmetrical three-phase winding (with phase shift between phases of 120\u00ba) separated by an angle (\u03b1 = 30\u00b0) Fig 2. The rotor is a squirrel cage type; the loops are shown in Fig 3. Multiple coupled circuit modeling for the DSIMs includes six differential voltage equations for the stator windings, Nb+1 differential voltage equations for the rotor meshes, and two mechanical differential equations. Thus, the model can be represented in the matrix form by : Vs1 = Rs1 \u00d7 Is1 + d dt \u03c8s1 (5) Vs2 = Rs2 \u00d7 Is2 + ddt \u03c8s2 (6) Vr = Rr \u00d7 Ir + d dt \u03c8r (7) Where: Vs1 = Vsa1 Vsb1 Vsc1 T (8) Vs2 = Vsa2 Vsb2 Vsc2 T (9) Is1 = Isa1 Isb1 Isc1 T (10) Is2 = Isa2 Isb2 Isc2 T (11) \u03c8s1 = \u03c8sa1 \u03c8sb1 \u03c8sc1 T (12) \u03c8s2 = \u03c8sa2 \u03c8sb2 \u03c8sc2 T (13) Vr = Vr1 Vr2 \u2026" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001763_aeat-04-2019-0087-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001763_aeat-04-2019-0087-Figure6-1.png", + "caption": "Figure 6 Edge controller prototype \u2013 close-up view", + "texts": [ + " To its left is the comparator chip and furthest away and to the left we spot the operational amplifier IC used to monitor the current by measuring the voltage drop across a current shunt. UAVs distributed avionics paradigm Svetoslav Zabunov and Roumen Nedkov Aircraft Engineering and Aerospace Technology Volume 92 \u00b7 Number 2 \u00b7 2020 \u00b7 229\u2013236 The six-pin connector is a prototype-only feature. This is the MCU programming and debugging interface. In the 3Dmodel and the PCB design schematic representation a female connector was used, but in the prototype boards, male connectors instead were used (see Figure 6). Last and furthest to the right on the board, one may see the MOSFET transistor managing the beacon light. A low noise and low dropout voltage regulator is responsible for the 5V stabilized power supply for the MCU and most of the on-board electronics. Mounting of the device is fulfilled by means of two 3mm bolts at two of the board edges as shown in Figures 5, 6 and 11. The version 8 prototype device can be observed in Figures 6, 7, 9 and 11. The difference from the 3Dmodel variant is the presence of an LED light beacon seen in Figure 6 at the edge of the PCB closest to the observer. Another distinction is the implementation of two instead of three tantalum capacitors (orange cubes further away in the blurred out-of-focus distant end of the board). The programming interface is relying on a male, instead of a female, connector, as mentioned already above. The connector is seen in front view having six golden pins. The acoustic altimeter is directlymounted to the controller. The motor cabling in red, yellow and black color is soldered at the appropriate board tracks and connects the BLDCmotor to the edge controller. The edge controller in Figure 6 is itself mounted to the fuselage of a small electrically powered UAV airplane. A larger view of the lattermay be seen inFigure 7. The controller is positioned at the middle between the empennage and the aircraft\u2019s main wing. Figure 7 shows the back section of the plane where the empennage with its two servomotors and the tail-mounted BLDCmotor with 12 8 in. pusher propeller are visible. The used motor is 200W, 700 KV type suitable for 10.8V three cell lithium-ion (Li-ion) battery pack. The airplane prototype has not been not tested so far; but, two other experimental setups using the edge controller were tested and results are disclosed later in the article" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003112_j.mechmachtheory.2020.103903-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003112_j.mechmachtheory.2020.103903-Figure4-1.png", + "caption": "Fig. 4. Hybrid serial-parallel 5-DOF 2-coupled-Cartesian-manipulator with parallel-revolute-axes, rotated \u221290 \u25e6 around \u02c6 Z W axis.", + "texts": [ + " Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx proven below, since revolute R joint axes \u02c6 Y Bn \u2016 \u02c6 Y Dn are geometrically parallel, there is no parasitic-twist-angle \u03b8BD n Z between links L Bn and L Dn so a passive revolute R joint is not required between them. Therefore links L Bn , L Dn , L T are rigidly connected, forming a single rigid body as shown in drawing In Sm Figs. 3 D, 4 D. The 5-DOF hybrid serial-parallel 2-coupled-Cartesianmanipulator with parallel-revolute-axes of Figs. 3 , 4 have joint topology ( PPP U )(P PP U ) . Note that they have four revolute R joints (two R \u2019s per U ) compared to the non-parallel-revolute-axes configuration Fig. 2 , with a total of five revolute R joints. The manipulator in Fig. 4 is identical to the one in Fig. 3 except that it is rotated \u03b8A 2 Z = \u221290 \u25e6, \u03b8C 2 Z = \u221290 \u25e6 around the \u02c6 Z W axis. The link and coordinate symbols in Fig. 4 are labeled with subscript \u2018 2 \u2019 to distinguish them from the ones in Fig. 3 with subscript \u2018 1 \u2019. Fig. 5 fully parallel 4-DOF 4-coupled-Cartesian-manipulator . Common-link L T in Fig. 5 connects the two 2-coupledCartesian-manipulators from Figs. 3 , 4 . Together they form the 4-DOF fully parallel 4-coupled-Cartesian-manipulator of Fig. 5 with 2 T 2 R motion-type. The passive revolute R joint, along the common-link \u02c6 Z T axis, accommodates parasitic-twistangle \u03b8BD n Z between links L B 1 , L D 1 and L B 2 , L D 2 since the revolute R joint axes of the two manipulators from Figs", + " The intersecting-revolute-axes version in Fig. 5 B, 5D has joint topology (P P UR )3( PP P U) , which may be represented as follow, suggesting the name \u2018Xactuator\u2019: Controlling the prismatic joint linear position z W A 2 in Fig. 5 B, 5D along vertical link L W adds position control z W T along \u02c6 Z W of the common-link L T for 5-DOF in a hybrid serial-parallel configuration, with joint topology ( P P P UR )3( PP P U) . In Fig. 5 , the 2-coupled-Cartesian-manipulator from Fig. 3 interleaves with the one from Fig. 4 so that link L B 2 from the second 2-coupled-Cartesian-manipulator is between links L B 1 , L D 1 from the first 2-coupled-Cartesian-manipulator. Similarly, link L D 1 from the first 2-coupled-Cartesian-manipulator is between links L B 2 , L D 2 from the second 2-coupled-Cartesianmanipulator. This allows coaxial bearings R 1 , R 2 to be spaced far apart from each other, along the common-link L T , as seen in Fig. 5 A, to support high moment loads with high moment stiffness. The load support link in solid model In Sm Fig", + " Mathematical description Kinematic equations [55] provide basis for the design, simulation, control and analysis of the coupled-Cartesian- manipulator family. The coordinate frames of Fig. 7 illustrate the 2-coupled-Cartesian-manipulators of Figs. 2\u20134 that are composed of two limbs, { L An , L Bn , L T } and { L Cn , L Dn , L T } respectively. The nomenclature in Fig. 7 pertains to generic links L An , L Bn , L Cn , L Dn where n = { 1 , 2 } but it applies specifically to links L A 1 , L B 1 , L C 1 , L D 1 in Fig. 3 or links L A 2 , L B 2 , L C 2 , L D 2 in Fig. 4 . The two 2-coupled manipulators from Figs. 3 , 4 , joined together by a revolute R joint along the common-link L T longitudinal \u02c6 Z T axis, form the 4-coupled-Cartesian-manipulators in Figs. 5 , 6 , each with four limbs. The general mathematical equations, derived for the 2-coupled manipulators in Figs. 3 , 4 , also apply to the 4-coupled manipulators in Figs. 5 , 6 . Coordinate frames. The positions and orientations of the components of the manipulators are expressed relative to Cartesian coordinate frames as shown in Fig", + " Similarly, if common-link L T connects to coupler-link L Dn then the orientation angles \u03b8C n X , \u03b8D n Y of links L Cn , L Dn depend on orientation angle \u03b8C n Z of link L Cn , \u03b8Cn X = atan 2 ( sin ( \u03b8Cn Z ) x W \u0302 Z T \u2212 cos ( \u03b8Cn Z ) y W \u0302 Z T , z W \u0302 Z T ) , 0 \u2264 \u03b8Cn X \u2264 2 \u03c0 (16) \u03b8Dn Y = sin \u22121 ( cos ( \u03b8Cn Z ) x W \u0302 Z T + sin ( \u03b8Cn Z ) y W \u0302 Z T ) , \u2212 \u03c0/ 2 \u2264 \u03b8Dn Y \u2264 \u03c0/ 2 (17) For the coupled-Cartesian-manipulators analyzed here, angles \u03b8A n Z , \u03b8C n Z in Eqs. (14 - 17 ) are fixed. For example, \u03b8A 1 Z = 0 \u25e6, \u03b8C 1 Z = 0 \u25e6 in Fig. 3 and \u03b8A 2 Z = \u221290 \u25e6, \u03b8C 2 Z = \u221290 \u25e6 in Fig. 4 . Inverse kinematics, common-link L T position T W . Given common-link L T desired position T W , the positions A n W , C n W of links L An , L Cn are derived from Eqs. (12 , 13 ) A n W = T W \u2212 R W B n ZXY T Bn \u2212 R W A n ZX B n An (18) C n W = T W \u2212 R W D n ZXY T Dn \u2212 R W C n ZX D n Cn (19) Given desired common-link L T position T W and orientation \u02c6 Z W T , all of the terms on the right hand sides of Eqs. (18 , 19 ) are known from equations (5 , 6 , 14-17 ) and the fixed geometry of the links B n An , T Bn , D n Cn , T Dn " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003733_s00158-020-02741-x-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003733_s00158-020-02741-x-Figure4-1.png", + "caption": "Fig. 4 The transition zone (arc)", + "texts": [ + " \u222bds1 \u00bc \u222b vdv g sin \u03b3\u2212\u03bcg cos \u03b3\u2212 CkA\u03c1av 2 2 m\u00feM\u00f0 \u00de \u00f06\u00de The integral was solved to obtain the relationship between the skiing velocity v1 at the end of the start zone and the related structural and environmental parameters. v1 ffiffiffiffiffiffiffiffiffiffiffiffiffi 1 Ck\u03c1aA s 1\u2212e CkA\u03c1as1 m\u00feM ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 m\u00feM\u00f0 \u00de g sin \u03b3\u2212\u03bcg cos \u03b3\u00f0 \u00de p \u00f07\u00de There were two kinds of profiles of the transition zone: the arc and third power function. The dynamic differential equations were established (Fig. 4). 1. Arc If the skiing velocity to a certain point of the inrun is v, then FN\u2212 m\u00feM\u00f0 \u00degsin\u03b8 \u00bc m\u00feM\u00f0 \u00de v 2 R ; \u00f08\u00de m\u00feM\u00f0 \u00deg cos\u2212F f \u2212 1 2 Ck\u03c1aAv 2 \u00bc m\u00feM\u00f0 \u00de dv dt : \u00f09\u00de Then, F f \u00bc \u03bcFN; v \u00bc R d\u03b8 dt ; m\u00feM\u00f0 \u00deg cos\u03b8\u2212\u03bc m\u00feM\u00f0 \u00degsin\u03b8\u00fe m\u00feM\u00f0 \u00de v 2 R \u2212 1 2 Ck\u03c1aAv 2 \u00bc m\u00feM\u00f0 \u00de dv dt \u00bc m\u00feM\u00f0 \u00de dv d\u03b8 v R ; dv d\u03b8 \u00bc gR cos\u03b8\u2212\u03bc sin \u03b8\u00f0 \u00de\u2212\u03bcv2\u2212Ck\u03c1aAv 2=2 m\u00feM\u00f0 \u00de v \u00f010\u00de where g is the gravitational acceleration (m/s2); R is the radius of the arc (m); \u03b8 is the central angle (\u00b0), \u03b8 = 90\u00b0-\u03c6; and \u03c6 is the inclination of the transition zone (\u00b0)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002956_022027-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002956_022027-Figure1-1.png", + "caption": "Figure 1. Relations between robot coordinate systems.", + "texts": [ + "1088/1742-6596/1550/2/022027 The matrix T0 6 represents the conversion relationship between the end link coordinate system and the base coordinate system, wherein the sub-matrix 0 6 N represents the attitude information of the end- link coordinate system in the base coordinate system of robot, and the sub-matrix 0 6 P represents the position information of the end-link coordinate system in the base coordinate system of robot. The tool installed on the end-link of the robot is the actuator of the robot. The tool coordinate system is the coordinate system fixed at the end of the tool. Its origin is the tool end center, called TCP point. The relationship among the base coordinate system, the end-link coordinate system and the tool coordinate system of the robot is shown in Figure 1. Through the kinematics model, it can be concluded that the representation of the end-link coordinate system E in the base coordinate system B can be represented by the transformation matrix T0 6 , while T6 tool is the transformation matrix of the tool coordinate system T relative to the end coordinate system E. If the effect of attitude is discarded, that is to say, the direction of tool coordinate system T and end coordinate system E is the same, and only the position relationship is different" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003977_icma49215.2020.9233548-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003977_icma49215.2020.9233548-Figure3-1.png", + "caption": "Fig. 3 Three kinds of virtual force action", + "texts": [], + "surrounding_texts": [ + "Lloyd coverage control is a method to cover area through Voronoi division. Suppose ip is the position coordinate of robot i . For a given area 2 iV , its mass density function is and the mass, centroid and polar moment of inertia about ip are ,, , i i i iV V V pM C J respectively, defined as [9] 2 , 1 ( ) , ( ) ( ) i i ii i i i i V V VV V V p i V M q dq C q q dq M J q p q dq (5) where q is any point in iV . Applying the parallel axis theorem, the polar moment of inertia about point pi can be written as 2 , ,i i i V i iiV p V C V i VJ J M p C (6) where ,i ViV CJ is defined as the polar moment of inertia of the area iV about its centroid. Setting 2 ( )i if q p q p , Eq. (2) can be transformed into 2 1 ( ) || || ( ) i N i i V H P q p q dq (7) Let ( ) ( )q q , we can derive another form of ( )H P and its partial derivative expression from (5), (6), (7) as, respectively, 2 , 1 1 ( ) i V i ii N N V C V i V i i H P J M p C (8) ( ) 2 ( ) i iV i V i H P M p C p (9) From the meaning of the gradient meaning of ( )H P , ( ) / 0iH P p can be used for minimizing the coverage performance function, that is, ii Vp C . In this paper, the first-order dynamic model and the control law (10) are used to control the movement of robot towards the center of mass [9], ( ) i i i i p i V p u u k p C (10) where pk is position gain and iu is the speed of ip . The robot is continuously moved to the centroid of its Voronoi partition through the control law (10). Simulation results given in Section V indicate the entire area cannot be covered completely. IV. IMPROVED LLOYD-VORONOI ALGORITHM BASED ON VIRTUAL FORCE The drawbacks in Lloyd [19], [20] are low coverage efficiency and incident perception interruption due to the existence of coverage holes for the boundary and the robot after deployment. In this context, we propose an improved Lloyd-Voronoi algorithm based on virtual force (LVVF) to settle the holes problem. After the Lloyd coverage, LVVF algorithm iteratively calculates the virtual force and drives robots to move by the virtual force. The algorithm stops when the coverage ratio meets the given condition." + ] + }, + { + "image_filename": "designv11_80_0001455_012165-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001455_012165-Figure5-1.png", + "caption": "Figure 5. Safety Factor ( Min 8.04 ul)", + "texts": [], + "surrounding_texts": [ + "IOP Conf. Series: Earth and Environmental Science 343 (2019) 012165 IOP Publishing doi:10.1088/1755-1315/343/1/012165" + ] + }, + { + "image_filename": "designv11_80_0002090_wzee48932.2019.8979999-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002090_wzee48932.2019.8979999-Figure1-1.png", + "caption": "Fig. 1 The investigated hydrogenerator region and FE mesh", + "texts": [ + " However, eddy currents in the stator and rotor laminations and the current skin effect in stator and rotor bars have been neglected. Rotor damper cage bars of the discussed hydrogenerator are short-circuited at the ends of individual poles; therefore, eddy currents induced in these rotor bars (with current density Jb) for a single pole fulfil the following condition: 0d 1 = = b b n b s b sJ , (3) where: nb \u2013 number of bars of damper cage placed in pole shoe, sb \u2013 cross-section area of a single bar of damper cage. The investigated region of hydrogenator\u2019s cross-section is shown in Fig. 1 together with FE mesh. The analysed area contains 69154 finite elements and 139359 nodes. In order to obtain unique solution of electromagnetic field equations (1-2), it is necessary to introduce proper boundary conditions. At the outer stator diameter \u03933, the zero Dirichlet boundary condition has been adopted for vector magnetic potential. The periodic conditions of magnetic potential have been set at the edges \u03931 and \u03932. Fig. 2 shows schemes of the external circuits of the stator and rotor windings connected to the field model of the hydrogenerator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure12-1.png", + "caption": "Figure 12. Displacement from a plain bearing to a textured surface (N = 6000rpm, W = 2000N)", + "texts": [], + "surrounding_texts": [ + "These\ufeff displacements\ufeff substantially\ufeff modify\ufeff the\ufeff radial\ufeff clearance\ufeff during\ufeff operation.\ufeff The\ufeff maximum\ufeff displacement\ufeffis\ufeffnoted\ufeffin\ufeffthe\ufeffangular\ufeffpositions\ufeffbetween\ufeff150\u00b0\ufeffand\ufeff200\u00b0.\nShear Stress Distribution Shear\ufeffstress\ufeffdistribution\ufeffin\ufeffmedian\ufeffplan\ufefftextured\ufeffplain\ufeffbearing\ufeffand\ufeffuntextured\ufeffplain\ufeffbearing\ufefffor\ufeffdifferent\ufeff radial\ufeffload,\ufeffis\ufeffpresented\ufeffin\ufeffFigure\ufeff15.\ufeffMaximum\ufeffshear\ufeffstress\ufeffis\ufeffrecorded\ufeffin\ufefftwo\ufeffangular\ufeffposition\ufeffof\ufeff the\ufeffplain\ufeffbearing\ufeff50\u00b0\ufeffto\ufeff70\u00b0\ufeffalso\ufeffat\ufeff220\u00b0\ufeffto\ufeff260\u00b0.\ufeffA\ufefftextured\ufeffplain\ufeffbearing\ufeffunder\ufeffa\ufeffload\ufeffof\ufeff10\ufeffkN\ufeffis\ufeff subjected\ufeffto\ufeffsignificant\ufeffshear\ufeffstress\ufeffcompared\ufeffto\ufeffa\ufeffnon-textured\ufeffplain\ufeffbearing.\nA\ufefftextured\ufeffplain\ufeffbearing\ufeffunder\ufeffa\ufeffload\ufeffof\ufeff10kN\ufeffand\ufeffa\ufeffrotational\ufeffspeed\ufeffof\ufeff6000rpm\ufeffis\ufeffsubjected\ufeffto\ufeff significant\ufeffshear\ufeffstress\ufeffcompared\ufeffto\ufeffa\ufeffnon-textured\ufeffbearing\ufeff(Figure\ufeff16).\ufeffThese\ufeffstresses\ufeffreach\ufeff40.105\ufeff Pa\ufefffor\ufeffa\ufefftextured\ufeffbearing\ufeffand\ufeffreach\ufeffonly\ufeff9,5.105\ufeffPa\ufefffor\ufeffa\ufeffnon-textured\ufeffplain\ufeffbearing.\nRotational Velocity Effect Pressure The\ufeffeffect\ufeffof\ufeffthe\ufeffrotational\ufeffspeed\ufeffof\ufeffthe\ufeffshaft\ufeffon\ufeffthe\ufeffpressure\ufeffdistribution\ufeffin\ufeffthe\ufeffplain\ufeffbearing\ufeffmedian\ufeff plane\ufeffof\ufeffthe.\ufeffThe\ufeffrotational\ufeffvelocity\ufeffof\ufeffthe\ufeffshaft\ufeffis\ufeffvaried\ufefffrom\ufeff2000\ufeffrpm\ufeffto\ufeff9000\ufeffrpm;\ufefffor\ufeffa\ufeffsupply\ufeff pressure\ufeffof\ufeffPa\ufeff=\ufeff0.04\ufeffMPa,\ufeffsupply\ufefftemperature\ufeffof\ufeffTa=\ufeff40\ufeff\u00b0C\ufeffand\ufeffa\ufeffradial\ufeffload\ufeffof\ufeffW\ufeff=\ufeff10000N.\ufeffThe\ufeff" + ] + }, + { + "image_filename": "designv11_80_0003144_s11665-020-04962-z-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003144_s11665-020-04962-z-Figure2-1.png", + "caption": "Fig. 2 Schematic of the shear test specimen after the bonding process", + "texts": [ + " The sample coding system in this study is as follows: TLP Temperature\u2014TLP Time. For example, for 1050-30 joint, the TLP temperature was 1050 C and the TLP time was 30 min. For microstructural examination, samples were produced from the TLP joints of size 10 mm 9 10 mm 9 10 mm using a wire cut machine. The TLP joints were characterized in detail using optical microscopy and scanning electron microscopy (SEM). For lap-shear testing, samples were produced from the TLP joints using wire cut machine as shown in Fig. 2. The overlap was 20 mm (four times the thickness of the specimen). Three tests were carried out at each condition on a universal testing machine and the average test result was reported. The shear test was carried out with the rate of 2 mm/min. Microhardness tests were performed across the TLP joints in various zones using a force of 0.5 N and with 10-15 s holding time. The fractured specimens were cut and prepared for SEM microstructural examination of the fractured surfaces. Figure 3 presents the microstructure of 1050-30 joint, which was observed by SEM in the secondary electron mode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002253_1350650120908116-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002253_1350650120908116-Figure3-1.png", + "caption": "Figure 3. Mesh and mismatched grid interfaces in the fluid volume.", + "texts": [ + " Numerical model description. In this part, a simulation model of the studied gear pump has been built. Based on the pump CAD modeling, the fluid volume has been imported into PumpLinx and then meshed. The computational domain includes inlet/outlet ports and pump internal volume itself. F2:5 h \u00bc Z 1 h s h 2:5 s\u00f0 \u00deds 0 h 4 4 exp 0:48349 1:72838 h 0:30963 h 2 4:3250 10 6 h 44 8< : \u00f011\u00de Correct mesh generation strategy of the flow passage can provide a satisfied grid template for the following simulation. Figure 3 shows the gird generation strategy of the gear engagement area. The computational domain is meshed by a body-fitted binary tree approach, which has a high degree of accuracy and efficiency. Furthermore, the gap between teeth is very small, so the mesh on that position is refined by six layers of the hexahedral element. A moving/sliding methodology is used to mesh the moving and static part separately, then these parts are connected by mismatched grid interfaces (MGI) technique. Flow factor solution strategy The procedure of flow factors calculation is shown in Figure 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001583_s00202-019-00889-4-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001583_s00202-019-00889-4-Figure9-1.png", + "caption": "Fig. 9 PM prototype machine. a Rotor and b stator (full view)", + "texts": [ + " Finally, a commutation vector is generated to obtain the needed power variation in each execution cycle of the algorithm. Figure\u00a08 shows the negative-saliency AFPMSG used in this work. The machine has a double rotor and central stator configuration [2]. (9) P\u2217 \u2212 \ud835\udf00P 2 > Pgen \u2192 P + + P\u2217 + \ud835\udf00P 2 < Pgen \u2192 P \u2212 \u2212 ( P\u2217 + \ud835\udf00P 2 \u2265 Pgen P\u2217 \u2212 \ud835\udf00P 2 \u2264 Pgen ) \u2192 P == Q\u2217 \u2212 \ud835\udf00Q 2 > Qgen \u2192 Q + + Q\u2217 + \ud835\udf00Q 2 < Qgen \u2192 Q \u2212 \u2212 The particular negative-saliency characteristic is achieved by replacing with soft iron a portion of the magnet material in the rotor poles (Fig.\u00a09a). Therefore, the d-axis inductance is increased, while q-axis inductance is almost not affected, leading to the condition that Ld is higher than Lq (negative saliency) corresponding to the inverse condition of typical PM machines [2]. The machine was designed, built and tested at the Electrical Engineering Department facilities of the University of Concepci\u00f3n, Chile. A full view of the machine is shown in Fig.\u00a09b. An experimental laboratory prototype has been implemented to experimentally verify of the proposed strategy. The system consists of a squirrel-cage induction machine (IM) that operates as a prime mover for the AFPMSG (Fig.\u00a010). The IM is supplied by a commercial variable frequency drive ABB ACS355. On the other hand, the generator is connected to a commercial drive Eurotherm 584S that has been properly modified to allow triggering the power devices with external control signals. Hall Effect sensors are used to measure the generator currents whereas a voltage transducer is used to measure the DC-link voltage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002901_012017-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002901_012017-Figure2-1.png", + "caption": "Figure 2. The decay of sprung and unsprung masses.", + "texts": [], + "surrounding_texts": [ + "It is important to study: maneuverability by following the speed of movement and lateral stability by minimizing the lateral slip angle. The four-wheel vehicle model was subsequently examined in the case of a low-level direction for ESP. The clutches were not controlled to a high level of control since two different vehicle models were considered. In addition, no lateral speed adjustment was provided, whereas a vehicle equipped with active front and rear steering systems (AFS and ARS) can allow lateral transient movement, for example to avoid an obstacle. The current state shows that we are using complex and reliable controllers based on unrelated simplified vehicle models, or that we are separating a complex vehicle model to use simplified controllers in each direction. The authors do not give priority to either the first or the second approach. The new approach has been sufficiently studied, in which a relatively sophisticated and robust high level controller is used, based on a relatively complex model of four-wheeled vehicle with an optimal coordination strategy. The objective is to evaluate dynamic coupling at the vehicle level to justify the structure of the high-level controller. A twin longitudinal transverse control requires a paired longitudinal transverse vehicle model. Future autonomous vehicles (AV) must simultaneously drive longitudinal, lateral and, ultimately, vertical controls. These speakers are connected because they are connected to the same system. For global chassis control (GCC), the vehicle\u2019s internal dynamic clutches must be considered. These connections may be more restrictive or, conversely, more relaxed depending on the chassis systems integrated in the same vehicle. For example, with two-wheel steering (2WS), a vehicle can avoid an obstacle only by changing its course (angle of movement). Unlike the vehicle 2WS, vehicle 4WS, four-wheel steering in one direction, you can bypass an obstacle without changing its direction. Since it is impossible to know the future hardware design of cars, and since the design of cars may differ from different manufacturers, the authors of the article are developing a new detailed global car simulation. In order to replace the driver and make cars autonomous, it is necessary to introduce additional embedded systems. In addition, in order to distinguish a car brand from competitors, various subsystems can be implemented by various manufacturers. One thing is certain, future vehicles can be heavily overloaded. A layered architecture is a good option for future automotive control tasks. This architecture provides, in particular, the following criteria: adaptability, fault tolerance, dynamic reconfiguration, extensibility and modularity. The car is equipped with an Active Rear Steering (ARS) system, a vehicle-based braking dynamics control system (VDC) and two rear wheel electric motors for vectorizing rear torque (RTV). The generalized efforts required to move the vehicle can be optimally distributed using optimization strategies based on optimization (CA) on four buses. Then, such bus forces can be converted into actuator commands and activate the corresponding system, avoiding any internal conflicts. DS ART 2019 IOP Conf. Series: Materials Science and Engineering 832 (2020) 012017 IOP Publishing doi:10.1088/1757-899X/832/1/012017 DS ART 2019 IOP Conf. Series: Materials Science and Engineering 832 (2020) 012017 IOP Publishing doi:10.1088/1757-899X/832/1/012017 There are three types of control: \u2022 High level control The goal here is to calculate the required forces on the vehicle propeller to track the desired speeds. Dynamics at this point is characterized by inertial parameters such as mass and moment of inertia. These parameters are usually uncertain, which requires a certain degree of reliability. A multi-input multi-output (MIMO) controller is required to account for various connections. Since the vehicle is equipped with ARS, VDC and RTV and access to active suspensions is not permitted, only a flat vehicle model can be considered at a high level. However, when evaluating tire potential, the importance of vertical dynamics should be considered. \u2211=Ss+Suf+Sur \u03a3: total body mass M and center of gravity (CoG) G, Ss: spring-loaded mass Ms and CoG Gs, Suf: front unsprung mass Muf and CoG Guf, Sur: rear unsprung mass Mur and CoG Gur. \u2022 Medium level control [8]. This intermediate level aims at coordinating the wheel systems in order to avoid differences and to generate high level general forces. To do this, all the forces must be distributed optimally on the four tires in order to activate the system with the desired level of effort. Since the number of forces on the tires exceeds the total forces that must be applied to the propeller of the vehicle, the system is overactivated, which leads us to the CA problem." + ] + }, + { + "image_filename": "designv11_80_0001739_ismsit.2019.8932756-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001739_ismsit.2019.8932756-Figure1-1.png", + "caption": "Fig. 1. Iteration 1", + "texts": [], + "surrounding_texts": [ + "978-1-7281-3789-6/19/$31.00 \u00a92019 IEEE\nKeywords\u2014 automation, gardening, mobile application, timely monitoring\nI. INTRODUCTION\nIn manufacturing sector, to control machine tool a processed is used which is called CNC machining. The process includes the use of computers to control the machine tools. Tools can be any object which helps us in performing different task like extruder in 3D printer, blade in lathe machine, drill bit in milling machine. CNC stands for Computer Numerical Control. It look like an ordinary personal computer (PC), but it differs due to unique software and console to control the machine.\nTo control the speed, location, coordinates and feed rate of machine a specialized CNC machining language is used, which is known as G-code. With the help of CNC machining we can control the position and velocity with great precision. CNC machining is currently being used in manufacturing for plastic and metal parts.\nThere are lots of advantages of using CNC machines. It\u2019s hard to manufacture any part in manual machining. Manual machining consumes more time and energy and it does not have much precision. While using CNC machining a lot of man power and time is secured and also the parts manufactured with CNC machining has great precision.\nA CNC machine consist multiple individual motors, which helps its tool to move in respective specific direction. A CNC Machine consist of two or more axis of motion. Usually CNC machine which are being used are consist 3-axis of motion. If there is any additional rotational movement then it will be 4-axis.\nII. LITERATURE REVIEW\nCommercially, Agriculture has been reached a newly high level of automation up till now, mainly for growing\ncrops on broad land. Fine-grain satellite imagery is also available commercially for pesticides and fertilizer related applications, leading to the precision agriculture\u2019s novel paradigm, its basic goal is to save water and pesticides [1]. The interest is growing day by day in autonomous farming so we can move towards smaller robotic platforms that can work on individual basis and can done some manipulation in the field with the help of precise sensing [2].\nWith the increasing research in the field of computer vision [3] and mechatronics which is helping us to led to an autonomous solution for harvesting some specialty crops which includes Cherries [4], apples, tomatoes, cucumber, mushrooms, strawberries, melons and much other.\nAnother active research on which work is being done is automatic weed control. Grey-level vision is used to navigate in structured outdoor environment and uses color vision to differentiate between the required crops and weed which is needed to be removed. Beside their multiple applications in the field, we envision automated agriculture robot which is precise and able to work without any operator in any sort of environment like urban areas, house roofs or can easily work in harsh environment like outer space.\nAlso, a lot of work is done in this field like autonomous targeted spraying [5] to removes pests from crops which can destroy and can have an effect on crops production, design of optimum manipulator for autonomous de-leafing [6] process of cucumber plant to prevent it from fungal disease which is considered by its growers but is costly. In addition to all of this there are also some virtual experimentation frameworks which have been developed for agriculture robots, [7], [8] ForboMind is a customized software platform which was introduced to support and help field robot done agriculture task done with precision and to promote to reuse the robotics components.\nOn the other hand, Agriculture field robots contributes to improve soil health, yield and reliability of operation. Which are commonly equipped with multiple sensors and cameras for navigation, localization, mapping and path planning algorithm.\nAlso, farming industry make use of drones for surveillance and monitoring fields. These drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides.", + "III. SOCIO-ECONOMIC SIGNIFICANCE\nAt present farming industry make use of drones for surveillance and monitoring fields these drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides. Similar Machinery exists but is only household specific; however we are aiming to target research and development departments in pharmaceutical and food industry, where setting up small greenhouse is required. Further we would be using some tool head for multiple operations instead of relying on magnetism for tool selection. Cutting down cost is another objective by using alternative materials to steel.\nIV. DESIGN\nWhile developing the project we went through some major and minor issues which lead us to make multiple iterations in the design phase. In this section we will discuss these iterations.\n Iteration 1:\nThis is the initial idea and concept on which we started to work. Our initial idea was to have x-axis, y-axis, z-axis and one rotary axis for the axis of motion. We also thought about adding multi-head mounter and add moisture sensor, shower and seeder. Initially we want to use rack and pinion mechanism. Issues: Rack and pinion mechanism can increase the vibrations which can affect the structure.\nWe then start to work on to select a mechanism for x-axis. We start to work on the stepper motor and lead screw mechanism. Issues: Our project has an open base so lead screw can\u2019t be mounted at the bottom like every CNC milling machine and usage of two stepper motor for just one axis can effect or destroy the structure if something happens to one motor.\n Iteration 3: So we started to think more about it and started to work on another solution to move the x-axis with just one motor. We finally decide to move the whole axis with one motor by using belt and pulley mechanism as shown in figure below.\nIn this project we use t-slotted extruded bars for developing our structure. Y-axis consist of simple lead screw mechanism.\nZ-axis consist of simple lead screw mechanism for linear motion. On z-axis a servo motor is mounted for giving a rotary motion to the multi head. The final design after the iterations is shown below.", + " Vacuum Pump: When a signal is sent out from controller to the vacuum pump and vacuum pump turn on and suck air and move towards seed when it come near the seed, then seed float and got stuck on the nozzle until the vacuum is on. When it reaches to the desired location then the controller sends another signal and vacuum pump turns off and the seed drops on the location the grid.\n Water Pump: When a signal is sent out from the controller to the water pump then water pump turn on and sprinkle the water over the grid with the help of shower for some specific time and then another signal comes from the controller and turns it off. Piping and instrumentation diagram (P&ID) is shown below:\nCurrently there are lots of global challenges which we are facing like global warming, food needs, poverty, degradation of climate and much more. Sustainable development goals introduced by United Nation helps us address these problems and base on these problems find solution to make future better not just for humans but also for the creature that exist on globe. The goals which can be achieved by our project \u201cGarden Tech\u201d are as follows:\n Zero Hunger: As we all know that human population is increasing rapidly to accommodate them deforestation is taking place. It is effecting our climate and results in global warming due to which glaciers are melting. Flooding is taking place more often than ever before and leads to devastation of arable land. Due to these circumstances food growth is not increasing rapidly like human population. The anticipated population according to United Nations Department of Economic and Social Affairs (UN DESA) is shown below:" + ] + }, + { + "image_filename": "designv11_80_0000276_ecai.2018.8679000-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000276_ecai.2018.8679000-Figure3-1.png", + "caption": "Fig. 3. Robot\u2019s pose as viewed from the target\u2019s position", + "texts": [ + " N significant observations (obs-1 to obs-N) have been constructed from the entire scanned data. These N observations are assigned 0 or 1 based on the presence of an obstacle in that region. A 0 value is allocated for the absence of any obstacle and 1 to indicate the presence of an obstacle. Some laser scanners have a relatively larger scan range in which case a sensitivity area needs to be defined. III. DILIGENT-BUG (D-BUG) ALGORITHM This new Diligent-Bug (D-Bug) algorithm primarily constitutes two parts: loop avoidance and selection of wallfollowing direction. Fig. 3 below indicates the orientation of a robot from the target\u2019s point of view. The definition of right, left and forward used in this paper is given as follows. Assuming that the laser scanner\u2019s coverage is 180o, \u03b8Ort indicates the orientation of the robot as seen from the target\u2019s point of view. When 0 \u03b8Rob \u03c0 2 and 3\u03c0 2 \u03b8Rob 2\u03c0 the target is assumed to be in front of the robot and hence it is forward. If \u03c0 2 < \u03b8Rob \u03c0, it is considered to be right while \u03c0< \u03b8Rob< 3\u03c0 2 indicates left. The Fig. 4 below indicates how a robot gets trapped in a Gshaped obstacle. There may also be several turns in the G-shaped obstacle. Whenever the robot reaches point A or B it is going forward thus getting stuck in an infinite loop. To avoid the robot from getting stuck in this kind of infinite loop, this D-Bug algorithm proposes the following. Consider the following Fig. 5 which indicates the forward (F), right (R) and left (L) poses of a given robot according to the definition given in Fig. 3 above. P = x y \u03b8Ort (1) R = cos\u03b8Ort sin\u03b8Ort 0 -sin\u03b8Ort cos\u03b8Ort 0 0 0 1 (2) The clockwise transitions of the robot\u2019s poses from F to L (FL), L to R (LR) and R to F (RF) are assigned a value 1 while anti-clockwise transitions from L to F (LF), R to L (RL) and F to R (FR) are assigned a value -1. Whenever the sum of these transitions becomes 0, the robot can resume its forward motion. Otherwise, it must stick to the left or right wall-following. It should be noted that if the initial pose of the robot is towards the target (forward position) or if the robot is repositioned to face towards the target before it begins its journey, the initial value of the accumulator needs to be set to zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003301_012032-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003301_012032-Figure3-1.png", + "caption": "Figure 3. Scheme of experimental research apparatus [7]", + "texts": [], + "surrounding_texts": [ + "This research was conducted using experimental methods. Tests carried out on the engine stand of a Kijang super 5K series engine car with a straight tube type radiator flat fin type. To obtain the effectiveness of the radiator whose engine rotation speed (low, middle and high) and also with time variations, then analyzed using the effectiveness equation (\u03b5) on the heat exchanger by measuring the temperature of the cooling water and cooling air. The process of installing the radiator and the whole series of tests was carried out at the Motor Fuel Laboratory and vehicle testing of the Department of Automotive Engineering, Faculty of Engineering, and State University of Padang Test equipment that will be used in this study include the following: Straight fin radiator, with specifications: Engine : Toyota Kijang 5K Fin model : straight fin Tube Model : flat tube Coolant type : water coolant Flow type : down flow Image of flat tube type radiator straight fin scheme on top view As for the research apparatus experiment scheme can be seen as shown below The 1st Progress in Science and Technology Research Symposium (PSTRS) 2019 Journal of Physics: Conference Series 1594 (2020) 012032 IOP Publishing doi:10.1088/1742-6596/1594/1/012032" + ] + }, + { + "image_filename": "designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.10-1.png", + "caption": "FIGURE 7.10", + "texts": [ + "14) that, for the microconvex surface, if the liquid drops are completely infiltrated into the solid surface, a and h remain constant, and with a decrease of b, the hydrophilic surface is more hydrophilic, and the hydrophobic surface is more hydrophobic. If b remains constant, with an increase of a and h, the hydrophilic surface becomes more hydrophilic, and the hydrophobic surface becomes more hydrophobic. It can be seen from Eq. (7.15) that for the microconvex surface, if there is a gas bearing between the solid liquid interface, its surface wettability is independent of h. If a remains constant, the surface hydrophobicity increases with the increase of b. If b remains constant, the surface hydrophobicity increases with the decrease of a. Fig. 7.10 shows the anisotropic-groove-textured surface. The width of the ridge between the groove is set to a (\u03bcm); the width of groove is b (\u03bcm); and the depth of groove is h (\u03bcm). For the Wenzel model, the roughness factor r is r5 11 h a1 b (7.16) Topography of anisotropic microgroove surface. For the Cassie Baxter model, the roughness factor f is as follows: f 5 a a1 b (7.17) Substitution of Eqs. (7.16) and (7.17) into Eqs. (7.9) and (7.11) gives cos\u03b8W 5 r \u03b3sv 2 \u03b3sl \u03b3lv 5 rUcos\u03b85 11 h a1 b\u00f0 \u00de cos\u03b8 (7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000530_978-3-030-12082-5_55-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000530_978-3-030-12082-5_55-Figure6-1.png", + "caption": "Fig. 6 Finite element mesh for nominal contact area (real contact area is darkened)", + "texts": [ + " The TCC parameter was set as a constant value equal to 157,000 W/(m2 K) or as the above-tabulated dependence on pressure TCC(p). The value 157,000 W/(m2 K) is obtained from the graph in Fig. 5 for the average pressure from the bolt clamp force of 7.3 MPa, calculated by dividing the sum of the clamp forces of each bolt (2000 N) by the nominal contact area. To evaluate the effect of thermal expansion, experiments were carried out for two types of contact behavior, Standard and No separation. For the Standard contact type, the contact heat transfer occurred strictly in the real contact area (Fig. 6), for which the TCC parameter was set. Thus, this type of contact reflects the influence of change in shape from thermal expansion. For the No separation contact type, movement of the contact surfaces along the contact plane is allowed, but separation of the surfaces is not permitted and the real contact area is equal to the nominal one. Thus, in the case of No separation contact, the change in shape of the surfaces from thermal expansion is not reflected in temperature results since it does not affect the thermal contact conductance", + " In this case, setting the dependence of the thermal contact conductance TCC on the contact pressure p obtained in the micromodel led to a decrease of 4.3 times in the averaged thermal contact conductance of themacromodel. The models used in the computational experiments Nos. 3 and 4 take into account the effect of thermal expansion on the real contact area. Because of the change in shape of the cylindrical body of the gyro unit, tangency takes place in the form of a narrow ring along the outer edge of the nominal contact area.Also, areas near the bolts are in direct contact. The real contact area was 56% of the nominal area (Fig. 6). As is clear from a comparison of experiments Nos. 3 and 1, the averaged thermal contact conductance decreased bymore than 5 times just due to accounting for the real contact area at a constant TCC of 157,000 W/(m2 K). Under the same conditions and using the TCC(p) dependence (experiments Nos. 2 and 4), the averaged thermal contact conductance decreased noticeably less, by 56%, which can be considered a result of thermal expansion without the direct influence of contact pressure. Repetition of the result of 56% is a random coincidence in this case" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure11-1.png", + "caption": "Figure 11 Fifth-order mode", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0003824_aiea51086.2020.00073-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003824_aiea51086.2020.00073-Figure3-1.png", + "caption": "Figure 3 Unclamped state diagram", + "texts": [ + " This often requires the preparation of multiple installation tools in the actual connection work In order to cope with the J-type clamps of different sizes and models, it brings a lot of inconvenience to the relevant operators, and also increases the design difficulty of the clamp installation tool[8]. III. STRUCTURE AND WORKING PRINCIPLE OF THE DEVICE The clamping device for installing tributary wires is a nonfixed position wire clamping method, with a large contact area and a fixed contact position, the clamping force is adjustable, and can be well adapted to different types of J components and different Diameter wires. The structure is shown in Figure 3 and Figure 4, including: Branch cable clamping arm, stop crossbar, fixed crossbar, moving crossbar, turning arm, rotating handle, connector, extension spring and other parts; 1- rotating handle; 2- turning arm; 3- stop crossbar; 4- branch cable clamping arm; 5- extension spring; 6- fixed crossbar; 7- moving crossbar; 8- connector The bottom end of the branch cable clamping arm is hinged to the shunt device, and the entire clamping arm can swing at a certain angle around the hinge point. The clamping arm is used to clamp the branch cable, so that the branch cable can be temporarily fixed in the lead groove of the J-type clamp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000815_iemdc.2019.8785385-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000815_iemdc.2019.8785385-Figure8-1.png", + "caption": "Fig. 8. Magnetic flux lines and magnetic flux density during no-load static analysis of a 3D-FEM simulation (a) and a 2D-FEM simulation (b). Note that the 3D model is rotated 45\u00b0 counterclockwise.", + "texts": [], + "surrounding_texts": [ + "To increase the accuracy in predicting the transient behavior during start-up of the unsymmetrical SPIM, the frequency dependent rotor resistance due to the skin effect in the rotor copper layer and the current dependent magnetizing inductance can be included for the d-axis and q-axis independently in the equivalent circuit model. More detail on non-linear effects is provided in the Appendix. Furthermore, improvements to the locked rotor test may be made, as assumptions generally made such as neglecting the magnetizing inductance LM seem to be invalid due to the large air gap of the machine under consideration. APPENDIX Fig. 9 illustrates the current dependent magnetizing inductance obtained by using the two different approaches introduced in IV-A. As expected, the magnetizing inductance depends on the saturation level of the machine, even more so in the auxiliary winding. Comparing the dashed and the solid red line, the leakage inductance of the auxiliary winding seems to be underestimated with a standard no-load test. Furthermore, the magnetizing inductance of the auxiliary winding saturates at lower current values when compared to the main winding saturation behavior. This can be explained by the saturating yoke due to the narrowing flux path of the auxiliary winding, Figs. 8(a) and 8(b)." + ] + }, + { + "image_filename": "designv11_80_0001924_icoecs46375.2019.8949914-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001924_icoecs46375.2019.8949914-Figure6-1.png", + "caption": "Fig. 6. The distribution of PM eddy current losses at a one phase short circuit (a) and in the absence of a short circuit (b)", + "texts": [], + "surrounding_texts": [ + "The paper presents an overview of the fault-tolerance technologies for EMs with PM for the transport application. It is shown that modern methods and designs can enable a fully fault-tolerant EM with PM. It is also shown that the use of 6-phase (and more phases) EM with PM is more effective than EM with duplex three phase system in terms of fault tolerance and efficiency. The use of bearingless solutions is one of the most promising technologies in the field of faulttolerant EM. Based on the analysis, an experimental prototype of fault-tolerant EM for FP was developed. To assess the bearings state of the experimental prototype, the diagnostic methods described above were used. In the future, it is planned to ensure full resiliency of this EM through the use of bearingless solutions." + ] + }, + { + "image_filename": "designv11_80_0003453_022004-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003453_022004-Figure2-1.png", + "caption": "Figure 2. Scheme of the disk working body.", + "texts": [ + " As a result, oil from spool valve housing 22 under pressure enters hydraulic cylinder 20, which leads to the retraction of rod 19, as a result of which hydraulic motor 7 with centrifugal disk 5 through lever 18 and rod 17 is rotated by an angle determined by the steepness of the slope. With decreasing steepness of the slope, the gauge wheel 31 with the associated pusher 30 under the influence of the spring 32 is lowered down. This, in turn, leads to a corresponding decrease in the angle of inclination of the hydraulic motor 7 and the centrifugal disk 5. The analysis of the results of the study of disk spreaders [10] shows that they require further improvement. The scheme of the proposed disk working body is shown in figure 2. Drum 1 with a radius rotates on wheels 2, in which rolling bearings are installed. Four droppers 3 with radius R are rigidly fixed on the drum. Let\u2019s consider the process of movement of a seed after it is separated from the ejector. A diagram of the forces acting on the seed is shown in figure 3. At any moment in time, when, 0900 i.e., when the seed falls from the ejector, the following equality holds: ,11 NP\u0420\u0421 += (1) where \u0421\u0420 is a centrifugal force; 1P and 1N are, respectively, the component weights of the seed and the reactions of the ejector, directed towards the axis of rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002334_j.matpr.2020.02.819-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002334_j.matpr.2020.02.819-Figure2-1.png", + "caption": "Fig. 2. Contact pair steel-PTFE", + "texts": [ + " The wagon structure is made so that there are three basic connection between its parts: connection between the tank and saddle (contact pair), middle connection (combination \u2013 contact pair and sliding bolt connection) and side connections between chassis of wagon and tank (combination \u2013 contact pair and sliding bolt connection; combination \u2013 contact pair and fixed bolt connection on Please cite this article as: M. Z\u030civkovic\u0301, V. Milovanovic\u0301, N. Jovanovic\u0301 et al., Experim for the transport petroleum products, Materials Today: Proceedings, https://do the side of hand brake). The tank is leaning on the saddle, which is in the direction of the wagon bogie, that is, the main transversal girders. For purpose relative sliding between the tank and saddle a layer of polytetrafluoroethylene (PTFE) is placed between them in order to reduce friction between steel plates (Fig. 2). Contact pair between tank and saddle is used as contact steel-PTFE with friction coefficient according [7]. As alreadymentioned, during the exploitation wagon is exposed to thermal loads. In order to allow free expansion and avoid occurrence of significant deformations, one side of the wagon is designed to allow relative sliding of the tank in relation to the wagon chassis. For this reason, the middle connection between tank and wagon chassis, as well as one side connection between tank and wagon chassis, is modelled with a sliding bolted connection, while on the other side connection between tank and wagon chassis is modelled with fixed bolted connection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003377_ccdc49329.2020.9164095-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003377_ccdc49329.2020.9164095-Figure4-1.png", + "caption": "Figure 4: Flight Configuration of SAW(\u03b4 = 60\u25e6)", + "texts": [], + "surrounding_texts": [ + "d\u03072k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)k(ld sin \u03b4k)\u03b4\u0307k \u2212(l cos \u03b4k)\u03b4\u0307k\n\u23a4 \u23a6 ,\nand\nd\u03082k\u22121(ld) =\n\u23a1 \u23a3\n0 0 0\n\u23a4 \u23a6 ,\nd\u03082k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)kld(cos \u03b4k \u03b4\u0307 2 k + sin \u03b4k \u03b4\u0308k)\nld(sin \u03b4k \u03b4\u0307 2 k \u2212 cos \u03b4k \u03b4\u0308k)\n\u23a4 \u23a6 .\nBecause the first derivative of the \u03b4k represents the folding angular rate, and in the assumptions presented above, the folding angular rate during morphing process keeps constant, so the second derivative of \u03b4k is zero, thus the second derivative of d2k can be expressed as follows:\nd\u03082k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)kld(cos \u03b4k)\u03b4\u0307 2 k\n(ld sin \u03b4k)\u03b4\u0307 2 k\n\u23a4 \u23a6 . (6)\nAfter computing d\u0307 and d\u0308 in equation (2) and (3) above, Meanwhile, the domain item dcm of Fext in (2) can also be decomposed in two parts:\nd\u0308cm = d\u0308cm1 + d\u0308cm2, (7)\nand\ndcm1 = 1\n2m mf lf (R1 +R3),\ndcm2 = 1\n2m mtipltip(R2 +R4),\n(8)\nwhere Ri(i = 1, 2, 3, 4) are represented as follows:\nR1 = 0, R2 = \u03b4\u030721 [0,\u2212 cos \u03b41, sin \u03b41] T , R3 = 0, R4 = \u03b4\u030722 [0, cos \u03b42, sin \u03b42] T ,\nSimilarly, the additional moment caused by morphing of wing tips can be computed. The domain item ( \u222b [d\u0304]d\u0308dm) in Mext can be given as follows: \u222b\n[d\u0304]d\u0308dm = 1\n2 [mf lf (S1 + S3)\n+ mtipltip(S2 + S4)][1, 0, 0] T , (9)\nwhere Si(i = 1, 2, 3, 4) represent the additional terms caused by folding of wing tip:\nS1 = 0, S2 = (l0 \u2212 ltip)\u03b4\u0307 2 1 sin \u03b41,\nS3 = 0, S4 = (ltip \u2212 l0)\u03b4\u0307 2 2 sin \u03b42,\nWhere l0 is the half width of the front view fuselage. Then, substituting the domain item d\u0308cm in Fext and\nthe domain item \u222b [d\u0304]d\u0308dm in Mext into the equation above, the Fext and Mext can be given as follows:\nFext = \u22121\n2 mtipltip(R\u03b41\u0394\u03b41 +R\u03b42\u0394\u03b42), (10)\nwhere R\u03b41 = \u03b4\u030721 [0, sin \u03b41, cos \u03b41] T ,\nR\u03b42 = \u03b4\u030722 [0,\u2212 sin \u03b42, cos \u03b42] T ,\nand\nMext = \u22121\n2 mtipltip(S\u03b41\u0394\u03b41 + S\u03b42\u0394\u03b42)[1, 0, 0]\nT , (11)\nwhere S\u03b41 = (l0 \u2212 ltip)\u03b4\u0307 2 1 cos \u03b41,\nS\u03b42 = (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42.\nThen, substituting equation (10) and (11) into equation (2) and (3), the nonlinear model of the folding wing-tip UAV can be expressed as follows:\nu\u0307 = rv + qw \u2212 g sin \u03b8 + Fx\nm , (12)\nv\u0307 = \u2212ur + wp+ gcos \u03b8sin\u03c6+ Fy\nm\n\u2212 mtipltip 2m (\u03b4\u030721 sin \u03b41\u0394\u03b41 \u2212 \u03b4\u030722 sin \u03b42\u0394\u03b42),\n(13)\nw\u0307 = uq \u2212 vp+ gcos \u03b8cos\u03c6+ Fz\nm\n\u2212 mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42),\n(14)\np\u0307 = (c0r + c1p)q + c2L\u0304+ c3N \u2212 c4 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42),\n(15)\nq\u0307 = c5pr \u2212 c6(p 2 \u2212 r2) + c7M, (16)\nr\u0307 = (c8p\u2212 c2r)q + c4L\u0304+ c9N, (17)\nwhere \u03b4\u03071,2 is the folding rate of the wing tip, \u0394\u03b41,2 is the folding angle change in a period of folding process, L\u0304,M ,N are components of total moment along three axes within body frame and can be given as follows:\nL\u0304 =p\u0307Ix \u2212 r\u0307Ixz + qr(Iz \u2212 Iy)\u2212 pqIxz \u2212 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42)\nM = Iy q\u0307 + pr(Ix \u2212 Iz) + (p2 \u2212 r2)Ixz\nN = r\u0307Iz \u2212 p\u0307Ixz + pq(Iy \u2212 Ix) + qrIxz\nwhere ci(i = 0, 1, 2, ..., 9) are constant coefficients expressed as follows:\nc0 = ( (Iy \u2212 Iz)Iz \u2212 I2xz\nIxIz \u2212 I2xz ), c1 = (Ix \u2212 Iy + Iz)Iz \u2212 Ixz IxIz \u2212 I2xz\nc2 = Iz\nIxIz \u2212 I2xz , c3 = Iz IxIz \u2212 I2xz , c4 = 1 IxIz \u2212 I2xz\nc5 = Iz \u2212 Ix\nIy , c6 = Ixz Iy , c7 = 1 Iy ,\nc8 = Ix(Ix \u2212 Iy) + I2xz\nIxIz \u2212 I2xz , c9 = Ix IxIz \u2212 I2xz\n1816 2020 Chinese Control And Decision Conference (CCDC 2020)\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply.", + "And Fx,Fy ,Fz are components of aerodynamic force and propulsive force along three axes within body frame, given that the proplusive force is along body x-axis and the engine offset angle \u03b1T = \u03b2T = 0, namely, T = Tx. Based on the conversion relationship between the body frame and airflow frame Xbody = ST \u03b1\u03b2Xwind, there exists:\n\u23a1 \u23a3 Fx\nFy Fz \u23a4 \u23a6 body = \u23a1 \u23a3 T 0 0 \u23a4 \u23a6 body + ST \u03b1\u03b2 \u23a1 \u23a3 \u2212D Y \u2212L \u23a4 \u23a6 wind (18)\nHence, Fx,Fy ,Fz are given as follows:\nFx=T + L sin\u03b1\u2212 Y cos\u03b1 sin\u03b2 \u2212D cos\u03b1 cos\u03b2, (19)\nFy = Y cos\u03b2 \u2212D sin\u03b2, (20)\nFz=\u2212L cos\u03b1\u2212 Y sin\u03b1 sin\u03b2 \u2212D sin\u03b1 cos\u03b2, (21)\nwhere L,Y ,D are thrust, side and drag force. \u03b1,\u03b2 are angleof-attack and the sideslip angle, \u03b8 and \u03c6 are pitch angle and roll angle.\nWith decoupling method proposed in [21], the decoupled longitudinal nonlinear model is given as follows:\nmV\u0307 = T cos\u03b1\u2212D+mg(\u2212cos\u03b1 sin \u03b8+sin\u03b1 cos \u03b8), (22)\nmV \u03bc\u0307 = T sin\u03b1+ L\u2212mg(sin\u03b1 sin \u03b8 + cos\u03b1 cos \u03b8)\n\u2212mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (23)\nmV \u03b1\u0307 = \u2212T sin\u03b1\u2212L+mV q+mg(sin\u03b1 sin \u03b8+cos\u03b1 cos \u03b8)\n\u2212mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (24)\n\u03b8\u0307 = q, (25)\nIy q\u0307 = M, (26)\nSimilarly, the decoupled lateral nonlinear model is given as follows:\nmV \u03b2\u0307 = Y \u2212mV (\u2212p sin\u03b1) + r cos\u03b1\n\u2212 mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (27)\n\u03c6\u0307 = p+ (r cos\u03c6+ q sin\u03c6) tan \u03b8, (28)\n\u03c8\u0307 = 1\ncos \u03b8 (r cos\u03c6+ q sin\u03c6), (29)\np\u0307 = (c0r + c1p)q + c2L\u0304+ c3N \u2212 c4 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42),\n(30)\nr\u0307 = (c8p\u2212 c2r)q + c4L\u0304+ c9N, (31)\nIt can be obtained from the equation (12)-(17) that during the morphing process of folding wing-tip UAV, some aerodynamic items such as the x-axis direction component of airspeed V , pitch and yaw angular rate, seem to have no distinction with that of conventional aircrafts. Consequently, It can be derived from above that the whole morphing process, including taking off, taking up and cruise, only has impacts on the aerodynamic performance of the folding wing-tip UAV in y-axis or z-axis, and in x-axis, the UAV appears the same as conventional aircrafts.\n3 Numerical Simulation\nTo validate the utility of the modeling method and nonlinear models of the folding wing-tip UAV, a numerical study is conducted to SAW, which is a typical kind of folding wing-tip UAV. The 3D model of the SAW is established in DATCOM and aerodynamic performances under different folding angles are analyzed based on numerical simulation. The basic airframe parameters of SAW are listed as follows:\nBased on above airframe parameters, with the application of DATCOM, the 3D model of the SAW can be established and the flight status(\u03b4 = \u221230\u25e6 and \u03b4 = 60\u25e6) can be shown as follows:\n2020 Chinese Control And Decision Conference (CCDC 2020) 1817\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply.", + "With DATCOM symbols and the corresponding shape parameter given above, the related parameters in DATCOM such as SSPNE and SSPN corresponding to different folding angles can be obtained and listed below, based on which numeric simulation can be conducted to obtain aerodynamic performance of SAW under different folding angles:\nThus, substituting all the parameters and symbol values into DATCOM, the aerodynamic coefficients with folding angles ranging from \u221230\u25e6 to 60\u25e6 as well as control surface items can be obtained. Based on this, the aerodynamic performances under different folding angles with angle of attack ranging from \u22124\u25e6 to 10\u25e6 can be shown below:\nIt can be illustrated in Figure 5 that during morphing process, the lift coefficients appear linear correlation with angle of attack, meanwhile, it appears that with folding angle keeping at \u221260\u25e6, the lift coefficients have a greater slope, which shows folding upwards can make it easier to enhance altitude when the SAW takes off. Similarly, in Figure 6, with folding angle keeps at 30\u25e6, the drag coefficients are much smaller, which make it more suitable in cruise phase.\nTo clearly clarify the ability of SAW to conduct multimissions, an insightful description of the polar curves of the SAW under different folding angles are given:\nIn Figure 8, when \u03b4 = 0\u25e6, the lift-to-drag ratio reaches the maximum value, which means that keeping the wing-tip level is more suitable to conduct long-range cruise surveillance missions. Folding wing-tip leads to the decrease of lift-to-drag ratio, which makes the SAW have the high-speed airfoil with small-aspect-ratio and large-sweep-angle. With the wing-tip folding upwards, the SAW can dive fast and conduct high-speed sprint, which has a greater maneuverability.\n1818 2020 Chinese Control And Decision Conference (CCDC 2020)\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure3.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure3.10-1.png", + "caption": "Figure 3.10 Diagram of forces acting on the element of the ring.", + "texts": [ + " The decrease of the bending moment when h/l = 1/8 and 1/6 is 13% and 22%, respectively, and the corresponding shear force equals to 14% and 13%, respectively. Thus, the conducted research showed that taking into account the shifting strain (deformation) when h/l\u2264 1/8 can lead to substantial changes in the maximum values of the reactions of a bent construction. 3.7 Oscillations of Circular Rings and Arches Let us examine flat bending (flexural) oscillations of circular rings following the work [298]. The diagram of the forces which act on the element of rings is shown in Figure 3.10. Differential equations of the angular, radial, and tangential motion of elements take the form (Figure 3.10) \ud835\udf15M \ud835\udf15\ud835\udf03 + \ud835\udc45\ud835\udc44 = \ud835\udefe\ud835\udc3c\ud835\udc45g \ud835\udf152\u03a6 \ud835\udf15t2 \ud835\udf15Q \ud835\udf15\ud835\udf03 + N = \ud835\udefe\ud835\udc39\ud835\udc45g \ud835\udf152U \ud835\udf15t2 (3.7.1) \ud835\udf15N \ud835\udf15\ud835\udf03 \u2212 Q = \ud835\udefe\ud835\udc39\ud835\udc45g \ud835\udf152W \ud835\udf15t2 M, Q, and N \u2013 bending moment, shear force, and the normal forces, acting in the section with the polar coordinate \ud835\udf03; F, I \u2013 area and the moment of inertia of transverse cross-section of the ring; \ud835\udefe \u2013 the volume weight of material of the ring; g \u2013 the acceleration of gravity; R \u2013 a radius of the ring; t \u2013 time; U, W \u2013 radial and tangential displacements; and \u03a6 \u2013 angle of rotation of the element of the ring with negligent of shifting strain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003457_02678292.2020.1813341-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003457_02678292.2020.1813341-Figure1-1.png", + "caption": "Figure 1. (Colour online) (a) Cylindrical coordinate system for director. (b) 3D cutaway view of the capillary (vertical view) and the ET director configuration.", + "texts": [ + " In this work, an ET configuration of nematic liquid crystals with chirality confined in cylindrical cavity is studied on the basis of liquid crystal elastic theory. We derive the equilibrium equations and the boundary equations under the planar anchoring conditions. Then two kinds of planar anchoring conditions which are azimuthal anchoring and axial anchoring respectively are investigated. Finally, the conclusions are given. A cylindrical coordinate system and a schematic diagram of ET director configuration are given in Figure 1. The director n which describes by the angle \u03b1 and the angle \u03b2, represents the statistical, average orientation of the liquid crystal molecules. \u03b1 is the angle between the director projection on r-\u03d5 plane and the er direction, and \u03b2 is the twist angle between the director n and ez axis. For cylindrical symmetry, \u03b1 and \u03b2 only depend on the variable r. We consider the director configuration of nematic liquid crystals in a cylindrical cavity with planar anchoring. The director n can be defined as: n \u00bc cos \u03b1 sin \u03b2er \u00fe sin \u03b1 sin \u03b2e\u03d5 \u00fe cos \u03b2ez: (1) The corresponding free energy density in the nematic director field is given by [32]: fel \u00bc 1 2 K11\u00bd \u00d1 \ufffd n\u00f0 \u00de 2 \u00fe K22 n \ufffd \u00d1\ufffd n\u00fe q0\u00f0 \u00de 2 \u00fe K33 n\ufffd \u00d1\ufffd n\u00f0 \u00de 2 K24\u00d1 \ufffd n\ufffd \u00d1\ufffd n\u00fe n\u00d1 \ufffd n\u00f0 \u00de\ufffd; (2) where q0 is the twist which is a chiral parameter of the chiral structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001500_iisa.2019.8900672-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001500_iisa.2019.8900672-Figure4-1.png", + "caption": "Fig. 4. (a) Prototype; (b) Design concept (top view) (figures from [17])", + "texts": [ + " Above a critical value of the rotational velocity of the minimotor, the actuation forces overcome the frictional forces at the contact points A and B, and the platform slides. Simulated results of the platform\u2019s trajectory are depicted in Fig. 3, where it is shown that during one cycle of operation, i.e. the eccentric mass has described an angle of 360\u25e6, the platform exhibits a net displacement in the Y -axis. The magnitude of the net displacement (step size) depends on the rotational velocity of the mini-motor [16]. 2) Dynamics of the mini-robot: An older prototype of the mini-robot is shown in Fig. 4a, see [17]. A new version of the mini-robot, based on the same actuation principle, is under construction. Some physical parameters of the mini-robot are presented in Table I. The dynamic model of the mini-robot is described by the Authorized licensed use limited to: Auckland University of Technology. Downloaded on May 28,2020 at 07:56:23 UTC from IEEE Xplore. Restrictions apply. following equations: Mv\u0307 = R \u2211 i bfi bI\u03c9\u0307p + b\u03c9p \u00d7 bIb\u03c9p = \u2211 i ( bri \u00d7 bfi ) + \u2211 j bnj (2) where i = {A,B,C,D,E}, and j = {D,E}, see Fig. 4b. In (2), M denotes the mass of the mini-robot, v = [x\u0307, y\u0307, z\u0307]> is the linear velocity of the center of mass of the mini-robot, and R is the rotation matrix from the body frame, {b}, to the inertial frame. bfi is a vector that includes the actuation forces generated by the two vibration motors and the friction forces at the three contact points of the mini-robot, and bnj includes the moments exerted by the vibration motors. The moment of inertia of the mini-robot is denoted by I, and \u03c9p is the angular velocity of the mini-robot. Finally, bri is the position vector of point i expressed in the body frame. The actuation forces generated by each vibration motor when its eccentric load rotates are given by the following equations: bfjX = (mr\u03b8\u0308 cos \u03b8 \u2212mr\u03b8\u03072 sin \u03b8) sin\u03c6j bfjZ = \u2212mg \u2212mr\u03b8\u0308 sin \u03b8 \u2212mr\u03b8\u03072 cos \u03b8 (3) where \u03c6j \u2208 {90\u25e6,\u221290\u25e6} is the angle between the motor axis and the X-axis of the body frame, see Fig. 4b. The Reinforcement Learning (RL) agent constitutes the basic building block of the proposed decision support system. The RL agent receives a state related to the position of the mini-robot and performs an action which corresponds to the direction of its velocity. Note that the desired magnitude of the mini-robots velocity is constant and not affected by the RL agent. The RL framework can be formally described as a Markov decision process (MDP) given by a five-tuple (S,A, T,R, \u03b3): \u2022 S denotes the set of agent\u2019s states" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000635_978-3-030-20377-1_1-Figure1.18-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000635_978-3-030-20377-1_1-Figure1.18-1.png", + "caption": "Fig. 1.18 Load history displayed in P-Q-\u03c3 space", + "texts": [ + " The chapter by Barber describes briefly more complex loading trajectories where the normal and shear forces form closed loops in P-Q space, but we will conclude with a brief description of the steady-state solution between two points 1, 2 in P-Q-\u03c3 space where the trajectory is in the form of a straight line. This is of some practical significance, because there are many cases of fretting contacts where a steady force develops a particular value of these quantities and then vibration, for example, introduces small changes in them all, so that P1 = Pmean \u2212 P/2 P2 = Pmean + P/2, (1.71) Q1 = Qmean \u2212 Q/2 Q2 = Qmean + Q/2, (1.72) \u03c31 = Pmean \u2212 \u03c3/2 \u03c32 = \u03c3mean + \u03c3/2. (1.73) It is easiest to think of the loading trajectory as a straight line between the end points, Fig. 1.18, but this might be relaxed to include some hysteresis provided that the slip zones advance monotonically during load changes in each direction. Because the normal load changes so will the size of the contact, and we also know that the slip zones will increase in size during each load change, and that as load reversals are experienced the entire contact will stick. Provided that the bulk tension effects are small, the shear traction in the slip regions will be the same at each end of the contact, and the maximum size of the slip zones will always be reached immediately before the reversal of load. Lastly, we know that the size and position of the stick zones at these extremes\u2014otherwise the \u2018permanent stick zone\u2019\u2014must be same so thatmaterial flows both in and out at each end of the contact andmaterial is preserved. Thus, the problem is solved when once we know a1 = a(P1), a2 = a(P2) together with the extent of the permanent stick zone [\u2212m n] which might, in principle, be a function of several quantities in Eqs. (1.72) and (1.73). Refer to Fig. 1.18 which is a sketch of the loading problem in P-Q-\u03c3 space, and Fig. 1.19 which shows the expected layout of the stick and slip zones as the end points 1, 2 are approached. When the load reverses the contact becomes fully stuck, and the slip zones grow monotonically as loading heads towards the other end of the trajectory, so that Fig. 1.19 shows the maximum extent of the slip zones and the minimum extent of the stick zone, i.e. the permanent stick zone. An important feature Fig. 1.19 Expected layout of the stick and slip zones as point 1, 2 is approached -a2 a2 -m a1-a1 n e of the sketch is that the permanent stick zone is of the same extent and at the same location at each end of the loading regime" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002873_icpes47639.2019.9105570-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002873_icpes47639.2019.9105570-Figure1-1.png", + "caption": "Fig. 1: Classification of different winding schemes by the number of slots per pole per phase q.", + "texts": [ + " Increasing the efficiency of the electric machine is preferable in all of these application fields, not only for environmental or economic reasons, but also in order to reduce the need for complex cooling systems. For higher rotor speeds, iron losses account for a significant part of the overall losses in the machine [2]. It is therefore desirable to have a fast estimation of the iron losses already during the design of the electric machine. Iron losses are highly dependent on the Magnetomotive Force (MMF) harmonics [3], [4], which can be influenced by the employed winding scheme. The different considered winding schemes can be classified by the number of slots per pole per phase (SPP) q, see Fig. 1. The concentrated fractional-slot windings (q < 1) produce a high number of harmonics and even sub-harmonics, especially for single-layer windings [5]. For integer-slot windings, the amplitudes of the MMF harmonics can be reduced The authors are with the Institute of Electrical Energy Conversion (iew), University of Stuttgart, Pfaffenwaldring 47, 70569 Stuttgart, Germany (e-mail: andreas.echle@iew.uni-stuttgart.de; urs.pecha@iew.uni-stuttgart.de; nejila.parspour@iew.uni-stuttgart.de). by increasing q (distributed windings) or by changing the ratio W /\u03c4p of the coil pitch W and the pole pitch \u03c4p (chorded windings)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001097_iciea.2019.8833640-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001097_iciea.2019.8833640-Figure3-1.png", + "caption": "Figure 3 \u2013 2-fingered gripper misalignment example.", + "texts": [ + " Although this binary criterion has been shown to be a useful measure to facilitate object gripping, it lacks any information related to how well the object has been handled \u2013 as evident by the fact that many methodologies [19, 20, 22, 27] are prone to a change in object orientation to some degree during the gripping process. The lack of descriptive power associated with this measurement \u2013 post-grip \u2013 is problematic for many industrial applications that require both controlled handling and placement of objects. In practice, a change in object orientation during the gripping process could be caused in part by a misalignment of gripper (figure 3) or gravity (figure 4). To better measure the affect the grasping process has on an object, a similarity metric is proposed in this paper to quantify the quality of candidate grasps. While the grasp only criterion is binary, a similarity score can be continuous. Such a score is calculated through a similarity test \u2013 in which an object is gripped and lifted vertically, moved away from the gripping location, moved back to its original location and placed at the gripping point. A new image is then captured and compared to its pre-gripped counterpart" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001404_pgsret.2019.8882707-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001404_pgsret.2019.8882707-Figure2-1.png", + "caption": "Fig. 2. Rotor slot structure and variables a) for lower cage, b) for upper cage structure of designed motor.", + "texts": [ + " the rotor cage, consists of aluminum and copper conductors, were analyzed with the same slot cross-sectional areas for equal comparison. As a result of the analyzes, the motor with aluminum conductors rotor is better about take-off performance [10,11]. In this study, the effects of changes in slot geometries on engine efficiency and take-off torque are analyzed to determine optimum rotor structure. The parameters representing the lower cage and upper cage of the double cage rotor are given in Figure 2. The slot structure given in Figure 2 was used in both the upper and lower cage of the rotor slots. Bs0 and Hs0 parameters, seen on Fig. 2(b), are representing the junction area of lower cage and upper cage. The variables that precise the lower cage and upper cage of the rotor and the initial values defined for these variables are seen on Table 2. Some of these variables were expressed as constant values considering design limitations The parametric solution method used in the optimization of electrical machines is a practical, fast and reliable method. In this method, the parameters given in Table 2 are defined as variables in the analytical solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure7-1.png", + "caption": "Figure 7. The solidification(a) is equality, compare with the formers. Defects(b)(c) happens in the gating system.", + "texts": [ + " Many defects located in the blades according to the combined defect parameter. Moreover, like the top gating system , the shrinkage defects and gas porosity stay in the similar place. The reason of defect is suspected to be the high speed of filling, which leading to a phenomena that not all the gas escape through the riser on time. ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 The results of linear-sharp side gating system is shown in Figure 7, which indicate that the quality of casting is satisfying and all the defects are inside the gating system that will not affect the impeller. The same thing also happens in cross-sharp side gating system. As the Figure 8. All the filling and solidification are stable, and effect-free defects in gating system and shrinkage defects in risers, the advance of side gating system is obvious. However, the solidification in cross-sharp gating system is more stable. As well as defect range in cross-sharp is smaller than the other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000182_012029-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000182_012029-Figure3-1.png", + "caption": "Figure 3. Installing the outer ring of bearing 126130 (d = 150 mm) in the housing with the thermocouples on the outer ring.", + "texts": [ + " To ensure a flow rate in range from 2 to 5 l\u00b7min-1, one nozzle (N1) was installed; for a flow rate in range from 5 to 7 l\u00b7min-1, two nozzles (N1 and N2) were installed; and for a flow rate up to 10 l\u00b7min-1, three nozzles (N1, N2 and N3) were installed. To determine the temperature state of the bearing, thermocouples were installed on the inner and outer rings of the bearing. Thermocouples were also installed to measure the oil temperature at the inlet and outlet of the bearing. The layout of the thermocouples is shown in figure 2. A photograph of the bearing with the installed thermocouples is shown in figure 3. ToPME IOP Conf. Series: Materials Science and Engineering 489 (2019) 012029 IOP Publishing doi:10.1088/1757-899X/489/1/012029 To plan an experimental study of heat generation in split inner-ring ball bearings it is necessary to choose the sufficient number of experiments to obtain an empirical relationship, by which it is possible to determine the heat generation in a ball bearing under different operational conditions. An analysis of theoretical and experimental researches of ball bearings [5, 6, 9], carried out earlier, showed that heat generation from friction in a ball bearing (Q) depends on the bore diameter of the bearing (d), rotational speed (n), axial load (Fa), oil flow rate through the bearing (V), and viscosity or oil temperature (T)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure23.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure23.3-1.png", + "caption": "Fig. 23.3 Arrangement of rod, coil, and tube a side and b front view", + "texts": [ + "1 shows the properties of the aluminum tube, copper oil, and steel rod used in this study. Simulations were carried out for different voltages levels using identical coil geometry which is modeled as a rigid body to make it simple for calculation in the boundary element method. Hence, deformation was analyzed only on the tube at the joining zone. The ranges of voltage were varied from 1.5 to 4 kV with an increasing rate of 0.5 kV. The arrangement of workpieces and tool coil for simulation is shown in Fig. 23.3. The model parts have meshed with 3290, 6580, and 2394 number of elements for the coil, rod, and tube, respectively. Electromagnetic crimping is a high-speed joining process using a magnetic pulse, and the material model which take care of strain rate hardening and softening due to temperature rise is Johnson cook (JC). In order to evaluate the constitutive response of the tube, the flow stress which is determined as a function of the plastic 274 G. T. Areda and S. D. Kore strain, strain rate, and temperature is related in JC constitutive equation as discussed by Schwer [8]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001552_s12555-019-0234-y-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001552_s12555-019-0234-y-Figure3-1.png", + "caption": "Fig. 3. Schematic of a 1-DOF planar discrete bending joint composed of N rolling units.", + "texts": [ + " 2 shows the fine positioning differences among the distal tips of three types of bending joints at the same bending angle: continuum bending, and pin- and rollingbased discrete bending joints. Assuming that the lengths of the central lines remain constant, between the rollingbased discrete bending joint and the pin-based discrete bending joint, the central line path and distal tip position of the rolling-based discrete bending joint are closer to those of the ideal continuum joint. Because each rolling unit is a combination of two adjacent identical pin joints, rolling-based bending results in a finer bending motion than pin-based bending with the same number of units [24]. Fig. 3 shows a uniform bending posture of the 1-DOF planar bending joint consisting of N rolling units each with an angle \u03b8u = \u03b8D/N, overall length H, and bending angle \u03b8D. The lengths of the proximal, intermediate, and distal links are Hprox, Hi, and Hdist , respectively, with the assumption that Hprox = Hdist = Hi/2 or H = NHi. R denotes the radius of the contacting curved rolling surfaces of the unit joints. Then, the homogeneous transform i\u22121 i T of the rolling units can be expressed as i\u22121 i T = cu \u2212su cu/2 ( 2R+(Hi \u22122R)cu/2 ) su cu 2Rsu/2 +(Hi \u22122R)su/2 0 0 1 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002957_022039-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002957_022039-Figure6-1.png", + "caption": "Figure 6. Robot repetitive movement", + "texts": [ + " For example, as depicted in Figure 5, if the input given was \u20181\u2019, then the robot shall move forward direction, if the robot received \u20182\u2019 or \u20183\u2019, it shall turn right and left respectively. By doing this, students will be able to observe the immediate effect of the if-else and switch structure in real life application. Module 3 emphasized on the use of repetition statement such as do-while, while, and for, to replicate certain robot action or movement pattern. The tasks include moving the robot in forward direction for two seconds, freeze for one second and repeat the action for three cycle recurrently as illustrated in Figure 6. During this module, students also taught on how to use suitable increment, decrement or assignment operators. In addition, the use of break and continue statements were also demonstrated to alter the repetition action of the robot. The fourth module utilize the concept of function to organize the programming statements and handling of I/O components. For instance, they learnt how to create numerous robot movement patterns and group the instructions within a set of function. Table 2 indicate the list of useable functions created to for the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000911_012043-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000911_012043-Figure1-1.png", + "caption": "Figure 1. Laboratory equipment (a), vacuum chamber with the table and wire-feeders (b) and the scheme of the EBAM process (c). 1 \u2013 Electron beam gun, 2 \u2013 electron beam, 3 \u2013 wire-feeder, 4 \u2013 wall, 5 \u2013 substrate, 6 \u2013 table", + "texts": [ + " However, the manufacture of such materials by the method of wire electron beam additive manufacturing (EBAM) has not been studied. Among other things, the production of multimetals based on titanium alloys, rather than technically pure titanium, also remains virtually unexplored. In this regard, the study of the formation of polymetallic metamaterials based on titanium alloys is a very urgent and promising task. Sample preparation was carried out using an experimental EBAM laboratory equipment at the Institute of Strength Physics and Materials Science SB RAS, Russia (Figure 1). Ti-6Al-4V alloy and AISI 321 stainless steel 1 mm diameter wires were used for EBAM deposition at process parameters shown in Table 1. PFSD-2019 IOP Conf. Series: Materials Science and Engineering 597 (2019) 012043 IOP Publishing doi:10.1088/1757-899X/597/1/012043 Templates were cut off the obtained sample, and then subjected to mechanical grinding, polishing with diamond paste and chemical etching in the reagent 1 (30 ml HNO3 + 0.5 ml HCl + 70 ml CH3COOH) during 30 seconds and in the reagent 2 (100 ml HNO3 + 1 ml HF) during 2 seconds" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001223_marss.2019.8860954-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001223_marss.2019.8860954-Figure4-1.png", + "caption": "Fig. 4: Experimental setup. In the inset image of agent, two spherical microrobots sit at the interface of water-air inside a glass petri dish. The identical agents shown have a radius of 400 \u00b5m. The agents are driven in horizontal plane by an electromagnetic coil system with three pairs of coils capable of producing fields in 3D.", + "texts": [ + " 2) Experimental setup: Magnetic fields for agent actuation are created in an electromagnetic coil system with three pairs of coils nested orthogonally to create fields in 3D, powered by three pairs of analog servo drives (30A8, Advanced Motion Controls). Each pair of wire loops in the coil system is arranged in Helmholtz configuration, resulting in a uniform magnetic field up to 15 mT (uniform to within 5% of nominal at the center over a workspace size of 5 cm) located at center of the coil system (see Fig. 4). The strength of magnetic field is smaller than the coercivity of the magnetic materials in the agents, and so the agents\u2019 magnetization will not be altered by the actuation field. Agent position is detected using a camera (FO134TC, FOculus) mounted atop the workspace, and a computer with custom Python code finds agent positions using a Hough Circle Transform in the OpenCV library at 60 frames/second. Two identical agents are immersed in a glass petri dish and sit at water-air interface as illustrated in Fig. 4. We tested the proposed method on two-agent configuration to evaluate the performance of the controller. Fig. 5 shows the experimental result for the controller to track a changing goal state. RMS tracking error of less than 140 micrometers and 5.68 degrees is accomplished for the regulation of the separation r and the pair heading angle \u03c6 , respectively. Thus, the frequency-based controller has the capability to operate as efficient as our previous controllers in [13]. Fig. 5 (a)-(d) presents four candidate snapshots associated to the given time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002906_pen.25409-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002906_pen.25409-Figure10-1.png", + "caption": "FIGURE 10 Predicted parison shape at time = 160 seconds, Weissenberg number Wi = 27.8 and for a die exit cross-sectional aspect ratio \u03c7 = 62.5, using a viscoelastic PTT model, A, extrudate shape, B, cross-sectional shape of the extruded sheet at a distance of 60 mm from the die exit showing thickness and width swelling", + "texts": [ + " In this case, the energy method developed previously (see Equation (38)) shows that the critical compressive stress for wrinkling increases drastically from \u03c3c = \u2212 1.317 MPa for \u03c7 = 12.5 to \u03c3c = \u2212 0.0524 MPa for \u03c7 = 62.5. In fact, the critical compressive stress for wrinkling scales like 1/\u03c72. In other words, an extruded sheet of very large aspect ratio provides lower critical stress and constraints for wrinkling. Let us study the shape of the extruded sheet at Weissenberg number, Wi, below critical value, for a fixed aspect ratio \u03c7 = 62.5. In Figure 10, we show the extrudate shape after 160 seconds of extrusion at Wi = 27.8 (corresponding to Vmax = 3.5 mm/s in Equation (39)), just below the critical Weissenberg number for the FIGURE 5 Predicted parison shape for a Newtonian fluid, A, without sag, B, with sag effect onset of wrinkling. In the same figure, we also show the cross-sectional shape of the extruded sheet at a distance of 60 mm from the die exit. The distribution of the first normal stress difference N1 = \u03c4zz \u2212 \u03c4yy along the central line axis, in the flow direction, is presented in Figure 9B. A minimum negative value of N1 = \u22120.0512 MPa is found just after the die exit. As this compressive stress induced in the extruded sheet soon after exit is high (N1 = \u22120.0512 MPa) compared to the critical compressive stress (\u03c3c = \u22120.0524 MPa) for the sheet to wrinkle, that is, N1 > \u03c3c, the extrudate remains flat (see Figure 10), in accordance with the energy method presented previously. Let us now study the effect of increasing the Weissenberg number close to the threshold of the instability, on the shape of the extruded sheet for a fixed aspect ratio \u03c7 = 62.5. In Figure 11, we show the extrudate shape at Wi = 29 (corresponding to Vmax = 4 mm/s in Equation (39)). Unlike the previous results at Wi = 27.8, the extruded sheet no longer remains flat. The crosssectional shape of the extruded sheet shown in Figure 11 deforms slightly out of the plane half its width" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000609_978-3-030-12684-1_17-Figure17.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000609_978-3-030-12684-1_17-Figure17.3-1.png", + "caption": "Fig. 17.3 Lattice ABS orientations and geometry", + "texts": [ + " The parts are designed to be rectangular cantilevers with high aspect ratio cross-sections to obtain relatively large displacements normal to the largest face and provide a wide, flat surface convenient for DIC speckling. Three solid ABS parts are printed with the dimensions shown in Fig. 17.1. Each part is built in a different orientation, as shown in Fig. 17.2. That is, all three parts in this set have the same outer geometry, but have their build layers oriented in different Cartesian directions. A second set of three ABS parts is printed with internal lattices. Each lattice is a 2D cell pattern printed along the build orientation, as shown in Fig. 17.3. Thus the second set of parts corresponds to the same build orientations as in Fig. 17.2, with lattices oriented vertically with respect to the base plate. The lattices are only built in the cantilever section of each part\u2014the base is left solid. The 2D cell pattern consists of 1.5 mm square holes separated by 1 mm solid walls, and surrounded by a 1 mm thick external wall. Figure 17.4 shows the outer dimensions of the lattice ABS set, which are increased from that of the solid ABS parts to allow for the internal lattice structures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001184_j.mechmachtheory.2019.103606-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001184_j.mechmachtheory.2019.103606-Figure9-1.png", + "caption": "Fig. 9. The inverted triangle chain and its equivalent limb.", + "texts": [ + " Therefore, it has been proved that the linear Delta mechanism can perform the three translational motions for the first time from the geometrical perspective. In addition, if the parallel mechanism is needed to achieve translational motion, the condition L 1 = L 3 must be guaranteed. 3.5. Equivalent mechanisms of two special parallel manipulators In all the 35 parallel manipulators with the same topology, the fourth and fifth mechanisms are quite special. For their limbs, there is a situation where the center of two spherical joints coincides, as shown in Fig. 9 (a). It is difficult to implement such a structure in practice. However, it should be noted that the upper beam and the two fixed length links form a triangular structure when the two spherical joints coincide. At this time, the two fixed length links connected between the moving platform and the slider is a stable structure. The 4S mechanism in the limb becomes the 3S one as shown in Fig. 9 (b). It should also be noted that the two spherical joints attached to the moving slider are not really free to move. The allowed movement is the rotation around the line formed by the two centers of the spherical joints. Thus, it is possible to replace the two spherical joints directly with one rotational pair. The spatial inverted triangle chain becomes to a PRS limb which everyone is very familiar with, as shown in Fig. 9 (c). The fifth combination ( C \u2207 C \u2207 C \u2207 ) would become the 3-PRS parallel manipulator as shown in Fig. 10 . Similarly, the spatial upright triangle chain can be equivalent to a PSR limb. The fourth combination ( C C C ) would become the 3-PSR parallel manipulator as shown in Fig. 11 . In the previous research, the established models of 3-PRS parallel mechanism are only suitable for the special structure (e.g., Fig. 10 (a)). For the more generic 3-PRS parallel manipulator shown in Fig. 10 (b), it is difficult to establish an analytical model because the coupling relationship between its output parameters has not been revealed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002889_phm-besancon49106.2020.00017-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002889_phm-besancon49106.2020.00017-Figure4-1.png", + "caption": "Fig. 4. Proposed crack propagation path", + "texts": [ + " During meshing process, there are two operating regimes, the first with one pair of teeth in contact and the second with two pairs of gears in contact. For one contact point, the next formula is used to determine the total stiffness: For the double contact period, the total stiffness is: A constant depth crack along the tooth width is considered in the gear tooth root and a quadratic curved path is proposed in this study for the crack propagation through the gear tooth thickness as illustrated in Fig. 4. The initiation point of the crack was determined by a stress analysis of the gear tooth subjected to the contact force. The presence of a crack causes a local reduction in the gear mesh stiffness due to gear tooth section area and moment of inertia decrease. The limiting lines proposed in this work for the tooth thickness reduction due to the presence of cracks with different lengths are represented in Fig. 4. These lines delimit the dead area corresponding to a specified crack length. Authorized licensed use limited to: University of Exeter. Downloaded on June 17,2020 at 04:06:54 UTC from IEEE Xplore. Restrictions apply. III. RESULTS AND DISCUSSION The gear parameters used for the simulation process are detailed in TABLE I. The natural frequencies of the undamped system are calculated using the mass and stiffness matrixes and . A mean value of the mesh stiffness is used and results are listed in TABLE II" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003955_ecce44975.2020.9235383-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003955_ecce44975.2020.9235383-Figure11-1.png", + "caption": "Fig. 11. Model of experimental electrostatic suspension apparatus", + "texts": [ + " 10 To determine the validity of the above analysis, and more generally, the feasibility of macro-scale electrostatic bearings, an experimental system has been designed. The system consists of a 130 mm diameter aluminum disk rotor, enclosed by two stator electrode arrays that are shown in Fig. 10. The stator outer and inner radii are dimensioned to match that of the rotor, to allow for equal charge distribution. The experimental conditions of the bearing are given in Table I. The experimental apparatus developed is shown in Fig. 11. To create the conductive electrodes, the stators of Fig. 10 take advantage of inexpensive printed circuit manufacturing and the rotor is simply a repurposed aluminum computer hard disk drive. Three laser displacement sensors, placed above the top stator, between the electrodes are used to determine the position of the rotor. Initial rotor air-gap and orientation is set using three fine adjustment screws. The whole apparatus will be placed in a translucent vacuum chamber to negate the breakdown limitation and allow for easy viewing of the levitation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001467_012017-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001467_012017-Figure2-1.png", + "caption": "Figure 2. Axis at Ordinate 1.25 at 2.08s, 6.66s, and 10 s", + "texts": [ + " Where in this simulation used the initial velocity of water on the river is 1.2 \ud835\udc5a/\ud835\udc60 and move faster based on time. The simulation will be shown in the picture below. ICETsAS 2018 Journal of Physics: Conference Series 1376 (2019) 012017 IOP Publishing doi:10.1088/1742-6596/1376/1/012017 From the simulation FIGURE 1, result at 1-meter axis ordinate, the maximum speed of water flow based on the simulation result is 34.6 \ud835\udc5a/\ud835\udc60. And the most stable flow is shown at 6.6 seconds. Based on the simulation FIGURE 2, result at the time of the ordinate of the shaft 1.25 meter, the maximum speed of water flow based on the simulation result is 28.4 \ud835\udc5a/\ud835\udc60. And the most stable flow shown at 6.6 seconds is also the same when the ordinate of the shaft is 1 meter. Based on the simulation FIGURE 3, result at the time of the ordinate of the shaft 1.5 meter, the maximum speed of water flow based on the simulation result is 26.1 \ud835\udc5a/\ud835\udc60. And the most stable flow shown at 6.6 seconds is also the same at the time of the ordinate of the shaft is 1 meter and 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002079_3352593.3352611-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002079_3352593.3352611-Figure2-1.png", + "caption": "Figure 2: Representation of a moving frame with respect to a fixed frame", + "texts": [ + " Hence, an interface to teach and learn individual transformations has been developed to solve this problem. Rigid-body motion in three-dimensional Cartesian space comprises of translation and rotation. The translation is defined using three Cartesian coordinates and the rotation needs three angular coordinates. Hence, the rigid-body motion can be defined completely using six coordinates. To identify the position and orientation of a body, a fixed reference coordinate system is used, which is called the fixed frame, represented using XYZ with origin O, in Figure 2. Another Cartesian coordinate system is attached to the moving body (frame UVW with origin at OM), to describe its position and orientation, with respect to the fixed coordinate system. The position of the moving frame (attached to a body) can be represented by a position vector from the origin of the fixed frame to the origin of the moving frame, \ud835\udc42\ud835\udc42\ud835\udc40 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7 , expressed in the reference of the fixed frame, as \ud835\udc42\ud835\udc42\ud835\udc40 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7\u20d7 \u20d7 = \ud835\udc42\ud835\udc40\ud835\udc65?\u0302? + \ud835\udc42\ud835\udc40\ud835\udc66\ud835\udc8b\u0302 + \ud835\udc42\ud835\udc40\ud835\udc67?\u0302? = [ \ud835\udc42\ud835\udc40\ud835\udc65 \ud835\udc42\ud835\udc40\ud835\udc66 \ud835\udc42\ud835\udc40\ud835\udc67 ] (1) However, to represent the orientation of the moving frame, the following important methods can be followed: 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000933_s10010-019-00352-7-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000933_s10010-019-00352-7-Figure1-1.png", + "caption": "Fig. 1 Drive principle of the CNC gear hobbing machine tool", + "texts": [ + " The improved ECCC architecture for five-axis linkage was proposed in Sect. 3. Besides, the characteristics of the ECCC architecture were also introduced. Sect. 4 illustrates the simulation experiments and the contour tracking property. Then all the experiments which are vital for proving the conclusion were performed on a 6-axis machine tool, the final results were analyzed in detail in Sect. 5. The final conclusions followed in Sect. 6. The six-axis gear hobbing CNC machine tool used in this paper is shown in Fig. 1. A is the hob angle adjustment axis, B is the hob spindle, C is the workpiece axis, X is the radial feeding axis, Y is the tangential feeding axis, and Z is the axial feeding axis. The synchronous motion of the workpiece rotational axis and the hob rotational axis is controlled by a master-slave EGB, implanted in the bottom of the gear hobbing CNC system. There are 3 axes linkages in using the axial-hobbing method to process a cylindrical helix gear: C-axis, B-axis and Z-axis. Prior to the movement of axes, EGB should be activated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003169_0142331220940484-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003169_0142331220940484-Figure7-1.png", + "caption": "Figure 7. The model of multiple-vertical-propellers aircraft.", + "texts": [ + " P= 0:1467 0:0334 0:0336 0:1518 ,F= 0:1282 0:028 0:0928 0:1253 : In system (29): u= 0:2635, and P= 0:2231 0:0301 0:0301 0:1985 ,F= 0:1461 0:0931 0:0931 0:1405 : In terms of the estimation of the domain of attraction, a comparison with these three kinds of systems is plotted in Figure 6. It can be verified that the estimation of the domain of attraction of single-layer saturated system is the largest one, which shows that there exists more conservatism with a bigger number of the layer of saturation. Example 3 Figure 7 is a model of multiple-vertical-propellers aircraft, which is employed to show the effectiveness of the presented controller design. We assume that there exist no air friction and transverse air drag. Here, we study the effect of x(t)= \u00bdc u f T to the fight of the aircraft under eventtriggered control. c is the yaw angel, u is the pitch angel and f is the roll angel. There exists nested saturation between the angels and the rate of propellers. A system model with the following form could describe the relationship between the parameters above and the flight displacement _x(t)=Ax(t)+B1sat(F1x(tk)+B2sat(F2x(tk))) S(t)=Cx(t) \u00f030\u00de where x(t)= \u00bdc u f T, S(t) is the displacement of the aircraft" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000630_s10704-019-00367-9-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000630_s10704-019-00367-9-Figure1-1.png", + "caption": "Fig. 1 Challenge test: AM 316L, 4mm thick plate subjected to uniaxial tension in Y-direction", + "texts": [ + "eywords Ductile tearing \u00b7 Additively manufactured metals \u00b7 Bifurcation \u00b7 Unified creep-plasticity The third Sandia Fracture Challenge (SFC3) issued by Sandia National Laboratories (Kramer et al. 2019) asked participants to generate blind crack initiation M. K. Neilsen (B) Sandia National Laboratories, Albuquerque, NM, USA e-mail: mkneils@sandia.gov and growth predictions for an additively manufactured 316L plate with a nominal gage section of 6.0mm width, 10.0mm height, with internal cavities subjected to uniaxial tension in the Y-direction (Fig. 1). Experiments were performed at a nominal displacement rate of 0.0126mm/s and analysts were asked to generate and report mechanical response predictions prior to seeing the experimental results. Ductile tearing predictions were generated using a Unified Creep Plasticity Damage (UCPD)Model with parameters for the model based on simulations of uniaxial and notched tension experiments. Numerous researchers (e.g. Tvergaard and Needleman 1984; Bammann et al. 1993; Xue et al. 2013) have investigated the simulation of ductile tearing with continuum damage models in finite element codes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003824_aiea51086.2020.00073-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003824_aiea51086.2020.00073-Figure6-1.png", + "caption": "FIG. 6 Clamping state", + "texts": [ + " When the turning arm swings around the hinge point, the connector on the turning arm pulls the spring, and the tension spring simultaneously pulls the branch cable clamping arm, and the branch cable clamping arm tightly bears the branch cable. With the turning arm continues to rotate, the contact point between the connector and the extension spring will pass the dead point formed by the turning arm, the extension spring, and the branch cable clamping arm, thereby ensuring The turning arm will not be pulled back to the starting point. After crossing the dead point, the turning arm leans steadily against the stop crossbar. As shown in Figure 6, the entire locking process is completed, and the branch cable is tightly locked in the wire groove of the clamp. So as to ensure the smooth completion of the entire branch cable installation. Finally, move the entire assembly to the stripping position of the main cable through a long insulating rod, start the motor to connect the main cable and the branch cable, and finally pull the separation wire to realize the separation of the clamping device and the J-type clamp, and the installation work is completed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003319_s10514-020-09938-5-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003319_s10514-020-09938-5-Figure1-1.png", + "caption": "Fig. 1 Orthosis mounted on the dummy", + "texts": [ + " The hip and knee joint angles are directly measured via magnetic angle sensors in the actuators, and velocities are measured using Hall-effect sensors in themotor and scaled by the transmission ratio. The ankle angle and velocity are inferred from the orientation and movement of the shank as measured by an inertial measurement unit (IMU) using sensor fusion of the gyroscope and accelerometer measurements via a manually-tuned complementary filter (Gui et al. 2015). The momentum-based control law described in this paper is experimentally evaluated on a prototype powered pediatric lower-limb orthosis frame attached to a test dummy, which is depicted in Fig. 1. The dummy allows free motion at the hip and knee joints and relies on bilateral ankle\u2013foot prostheses for passive stiffness and damping at the ankle joints. The components are sized for the body segment lengths of an average 8 year old child (Fryar et al. 2012; Winter 2009), with link lengths tabulated in Table 1. In the proceeding subsections, the equation of motion is described for the legs of the system, a ZMP analysis is discussed, and the system model is identified. For the purpose of identification of the complete orthosis-dummy system, the triple pendulummodel is identified separately from the ankle stiffnessmodel since linear regression techniques are directly applicable to robotic systems such as pendulums but are not suitable for the ankle model identification" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002435_012137-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002435_012137-Figure1-1.png", + "caption": "Figure 1. Permanent Magnet DC Motor Circuit", + "texts": [ + " PSpice has also been used for simulation of electro-mechanical devices, such as induction motor in industrial application, as written in [5]. LTSpice has been used widely in electrical and electronic students around the world, such as the following report written in [6]. Permanent Magnet DC Motor (PMDC) is a class of motor in which the stator has a fixed field, due to the permanent magnet attached to it. This type of motor has a linear characteristic with regards to its excitation voltage at the rotor. And therefore, the control of speed can be done, by adjusting the voltage terminal at the rotor field. Based on Figure 1, the following formulas are used: \ud835\udc63 = \ud835\udc45\ud835\udc56 + \ud835\udc3f \ud835\udc51\ud835\udc56 \ud835\udc51\ud835\udc61 + \ud835\udc49\ud835\udc52 (1) The 3rd International Conference on Eco Engineering Development IOP Conf. Series: Earth and Environmental Science 426 (2020) 012137 IOP Publishing doi:10.1088/1755-1315/426/1/012137 In order to understand the basic principle of DC Motor, the model of DC Motor is simplified to contain only the resistance of its coil, as shown in Figure 2. The circuit diagram in Figure 2, model the DC Motor as it is only consisting of pure resistance only" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003043_jsen.2020.3007503-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003043_jsen.2020.3007503-Figure1-1.png", + "caption": "Fig. 1. Two-dimensional serpentine flexible manipulator.", + "texts": [ + " At last, Section V concludes this paper. Compared with continuum flexible manipulator, serpentine flexible manipulator has advantage of high load capacity, which has great potential for medical applications [20]. And two-dimensional serpentine flexible manipulator can avoid the twist of the manipulator. In this section, the kinematic model and statics model of a two-dimensional serpentine flexible manipulator are introduced. The structure of the joint of the serpentine flexible manipulator is shown in Figure 1, which only allows the motion in the plane. The vertebras are serially linked, with adjacent vertebras forming a revolute joint. It can bend side to side in a plane. The joint motion is confined by the elastic component (an elastic tube). And the 3D motion of flexible manipulator can be realized by rotating the base of the flexible manipulator. The serpentine flexible manipulator has N vertebras and is driven by two wires, as shown in Figure 1. 1T and 2T are the tensile forces of two wires; 0h is the distance between each two vertebras; j stands for the bending angle of the j-th vertebra; H is the length of vertebra. The position information of i-th vertebra ( ), ( )x i y i can be expressed: 0 1 1 0 1 1 ( ) ( )cos ( ) ( )sin N i j i j N i j i j x i H h y i H h = = = = = + = + (1) The wires\u2019 length change can be calculated by (2). 1L and 2L are the length change of the two wires; d is the distance between the two wires", + " F is the foundation matrix which can be obtained by planar pattern calibration. From reference [7], we can know that Bezier curve can be well used to fit the shape of flexible manipulators. And the Bezier curve can be expressed: 3 2 2 3 0 1 2 3( ) (1 ) 3(1 ) 3(1 ) 3t t t t t t t= \u2212 + \u2212 + \u2212 +B P P P P (8) where 0P and 3P are the start point and the end point of a cubic Bezier curve respectively. 1P and 2P are the control points. When /t i N= , ( / )i NB is the coordinates of the i-th joint. As shown in Figure 1, point 0P and point 1P are on the same line. Therefore, the point 1P can be expressed: 1 0 01 0P P D= + S (9) where 0 0 1 01/D P P= S and 01 0 1S = P P . We assume that the position and direction of the first vertebra are given, which can be easily obtained by several measuring instruments, such as angle encoder and raster. Thus, the known parameters are ( 0P , 0D , 1T \uff0c 2T \uff0c 1L \uff0c 2L , bQ ). The length change ( 1L \uff0c 2L ) and tensile forces ( 1T \uff0c 2T ) of the wires can be obtained by encoders and force sensors, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure4-1.png", + "caption": "Figure 4. Meshing of model with top gating system.", + "texts": [ + " Then the model that including four types of gating system are created. Top gating system, bottom gating system and 2 types of side gating system are added to the impeller as Figure 3. ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 2.2.Simulation of gating system The software Anycasting bears the simulation. The mesh are created in the preprocessing module called AnyPRE. Taking the top gating system as the example, the account of flexible finite elements is 103488 in Figure 4. Setting the case as the filling analysis and solidification analysis, then the results of preprocessing are saved for later treatments. Solver module is used for calculating the results, as the results will be displayed as image by post-processing module. More settings are as bellowed Table 2. Combined defected parameter and probabilistic defect parameter are two evidences that for determining the filling results[7]. ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001428_012115-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001428_012115-Figure1-1.png", + "caption": "Figure 1. Scheme of the hardening fasteners billet process by transverse running with smooth plates", + "texts": [ + " Therefore, it became necessary to study this process in depth, to develop a general engineering methodology, which allows not only to predict the optimal choice of the processing modes main parameters, but also to carry out these parameters\u2019 targeted management in the process of their practical implementation. The purpose of this work is to determine the condition of gripping the workpiece, the stress state and the quality of the fasteners\u2019 surface layer during with flat plates. Scheme of the transverse running process. Consider the scheme of the process of hardening fasteners billet by transverse running with smooth plates. Figure 1 shows that the geometrical model of the running consists of the bottom rigidly fixed plate 3 and the upper moving plate 1 moving in the horizontal direction with the speed V. Between the plates placed blank 2 with a diameter of D. The main transverse running parameters are the absolute compression and tool geometry. In the transverse running process, the tool has the appearance of a flat plate with a small angle of the inlet part \u03b11 (Fig. 1). A small angle \u03b12 in the output part of the tool serves to reduce the stress concentration when the part leaves the machining area. To implement this process of running it is necessary to determine its main parameters: the capture angle of the workpiece, absolute compression, stress state in the deformation zone, residual stresses after running. To carry out the process of running the blank with flat plates, it is necessary to create certain conditions. Moreover, it is necessary to consider separately the conditions for the unsteady running process \u2013 for ICI2AE 2019 IOP Conf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002432_humanoids43949.2019.9035008-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002432_humanoids43949.2019.9035008-Figure4-1.png", + "caption": "Fig. 4. Skeletal structure of Kengoro\u2019s leg and force sensor configuration on its foot.", + "texts": [ + " Toe joint is effective to make humanlike natural postures and motions of robots, because toe DOF helps to solve inverse kinematics during knee extended posture of humanoids. Step motion can be generated from kneestretched standing posture although knee-bended squatting posture is initial posture of ordinary humanoids. The feet of Kengoro are equipped with multiple single axis force sensors (load cells) on its soles [12]. As the location of each sensor, 5 is at each tip toe, 5 is at the base of finger and 2 is at the heel. The overview of the foot and its sensor configuration are shown in Fig.4. The position of zero moment point (ZMP) [13], p, is calculated with the following equation. Where pj(pjx, pjy) is position of each load cell, and fjz is each force measured from the load cells. The number N of load cells in each foot is 12. We use ZMP as the indicator for ankle-hip stabilizer. px = \u2211N j=1 pjxfjz\u2211N j=1 fjz , py = \u2211N j=1 pjyfjz\u2211N j=1 fjz (1) p = [px, py, pz] T (2) III. IMPLEMENTATION OF ANKLE-HIP-STEPPING STABILIZER Fig.5 shows the flowchart of ankle-hip-stepping stabilizer which is implemented as integration of ankle-hip and stepping stabilizers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003672_s42835-020-00538-y-Figure16-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003672_s42835-020-00538-y-Figure16-1.png", + "caption": "Fig. 16 Prototype of the PMSpM. a After completing assembly, b rotor shell, c stepped PM, d stator shell, e support structure, and f stator coil", + "texts": [ + " The opening above the rotor is used to install the output shaft and stepped PMs are installed in the opening of the rotor structure. The stator is divided into two layers, which are connected by coupling bolts. The stator shell is made of resin material in order to reduce the weight of the entire motor. There are seven support structures on the lower stator shell to ensure the stable motion of the rotor. The coil is installed in the mounting hole of the stator structure. The prototype of the PMSpM is shown in Fig.\u00a016. The PMSpM has 3-D motion characteristics, and traditional torque measurement methods are no longer applicable [28]. The research group measures the acceleration of the rotor at the start of the motion through a Micro-ElectroMechanical System (MEMS) gyroscope sensor and calculates the torque of the motor by combining the motor kinematics equation [20, 29, 30]. The MEMS sensor is Mpu6050, which is installed on the top of the motor output shaft. The installation structure integrates the Bluetooth data-transmission module and the chip power-supply module" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003150_csei50228.2020.9142512-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003150_csei50228.2020.9142512-Figure6-1.png", + "caption": "Figure 6. The interface of the gear drive experiment", + "texts": [ + " When students take a mechanical design course, they can use the virtual remote lab to view the corresponding equipment for course preparation or review, and a step by step from easy to a complex scheme (See Figure5.) is used in building the online experiments. Some experiments interfaces are listed to better illustrate this scheme. For example, when learning drive mechanisms, the virtual lab enables students to operate fundamental experiment at the beginning, after finishing a simple one, a more complex experiment will come, this easy to difficult (see Figure6.) step by step experiments scheme in each individual chapter can consolidate students\u2019 comprehension systematically. a) The interface of the helical gear experiment Also, this step by step scheme is adopted in the whole teaching content. After the learning of each chapter, learners will have a comprehensive and global understanding of the mechanical design course, and to better strengthen the knowledge point and connect the theory of each chapter to practice, several engineering applications are also added at the final part of the course" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000609_978-3-030-12684-1_17-Figure17.7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000609_978-3-030-12684-1_17-Figure17.7-1.png", + "caption": "Fig. 17.7 Mode classes found include (a) out of plane (OOP), (b) torsion (TOR), (c) in plane (IP), and (d) extension (EXT)", + "texts": [ + " The maximum average amplitude was typically about 0.1 mm for the lowest mode, against a noise floor in the frequency domain of 30 nm. A baseline finite element model is developed using ABAQUS Finite Element software. Mesh fidelity along the cantilever cross section was varied between a shell model, a solid model two elements thick, and a solid model four elements thick. Modes are placed into four distinct classes, out of plane (OOP), torsion (TOR), in plane (IP) and extension (EXT), as shown in Fig. 17.7. The modal frequencies of the first four OOP modes converge at the solid model two elements thick, justifying the selection of this mesh for subsequent simulations (Table 17.1). Quadratic elements (C3D20R) are used with 2 \u00d7 20 elements for the cantilever cross section and 70 elements along the cantilever length. The inner faces of all four bolt holes are assumed to be fixed. Material properties are defined by an elastic orthotropic material model, presented in Eq. (17.2). The \u03b512 equation has analogues in the 13 and 23 planes, making the matrix symmetric such that only 3 of the 6 Poisson\u2019s ratios are independent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003328_012034-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003328_012034-Figure1-1.png", + "caption": "Figure 1. Loads on the power screw.", + "texts": [ + " For greater accuracy with the load data that the shaft will support, the weight of the weld deposition was calculated taking into account the ER70S-6 welding feed material for a linear bead of 15 cm, and a welding speed of 1.44 mm/s. In addition, the weight of the plates was added to the calculations. The physical properties are shown in Table 1. On the other hand, to avoid damage to the linear table, either by current leakage or by excessive heating during the welding process, a Bakelite plate was necessary, a copper plate to facilitate rapid heat distribution and return of welding current, this allows electrical and thermal insulation, the parts are shown in Figure 1. 6th International Week of Science, Technology and Innovation (6th IWSTI) Journal of Physics: Conference Series 1587 (2020) 012034 IOP Publishing doi:10.1088/1742-6596/1587/1/012034 Power screws are mechanical devices that change a rotation or angular displacement in a straight line, transmitting force and mechanical power. In practice, power screws are provided by specialized supplies that offer technical literature that includes all the data necessary for their selection. Two common types of thread are shown in Figure 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001303_chicc.2019.8865927-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001303_chicc.2019.8865927-Figure2-1.png", + "caption": "Fig. 2 Phase structure of multi-UAV cooperative tracking", + "texts": [ + " In order to guarantee the tracking effect, that is, to ensure that the intelligent target is within the observation field of at least one UAV at any time, the formation control tracking strategy centering on the target is introduced. The ground target is regarded as the centroid of formation structure. That is to say, the formation control strategy does not aim to construct formation among UAVs for target tracking, but to regard ground targets as part of formation structure itself and maintain tracking status by controlling formation structure. As shown in Fig. 2, three UAVs and a ground target form a target-centered 4-unit formation structure in a 2D plane. The main control quantity is the distance coordination between each UAV and the target, as well as the phase relationship of each UAV. Essentially, based on a leader-follower formation system, this structure regards the ground maneuvering target as leader and the operational UAVs as followers. In order to ensure that the target is always kept in the formation system and can be monitored by at least one follower, the distance between the leader and followers and the phase relationship between followers are adjusted through the coordination of position and speed between followers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003453_022004-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003453_022004-Figure3-1.png", + "caption": "Figure 3. Scheme of forces acting on the seed.", + "texts": [ + " This, in turn, leads to a corresponding decrease in the angle of inclination of the hydraulic motor 7 and the centrifugal disk 5. The analysis of the results of the study of disk spreaders [10] shows that they require further improvement. The scheme of the proposed disk working body is shown in figure 2. Drum 1 with a radius rotates on wheels 2, in which rolling bearings are installed. Four droppers 3 with radius R are rigidly fixed on the drum. Let\u2019s consider the process of movement of a seed after it is separated from the ejector. A diagram of the forces acting on the seed is shown in figure 3. At any moment in time, when, 0900 i.e., when the seed falls from the ejector, the following equality holds: ,11 NP\u0420\u0421 += (1) where \u0421\u0420 is a centrifugal force; 1P and 1N are, respectively, the component weights of the seed and the reactions of the ejector, directed towards the axis of rotation. The condition for separation of the seed from the ejector is 01 =N . Thus, at the time of separation of the seed equality 1P\u0420\u0421 = will be true, or, cos2 mgm = ,whence: .cos 2 g = (2) The formula (2) gives the relationship between the polar coordinates of the seed at the time of separation from the ejector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002532_ropec48299.2019.9057139-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002532_ropec48299.2019.9057139-Figure3-1.png", + "caption": "Fig. 3. Different types of bearing damage. a) Drilling on an external track 3.1 mm in diameter. b) Drilling on an external track 1.58 mm in diameter. c) Damage to the ballast by drilling through the cage, partially drilling the ballast to a diameter of 1.58 mm in diameter. d) Distributed damage due to corrosion and wear.", + "texts": [ + " Damage to the ball is done by drilling a ball through the cage, with a diameter of 1.6 mm, the fourth type of damage consists of a bearing exposed to corrosion, extracted at the end of its useful life from an engine that was used in the industry; therefore it is considered a distributed damage. The damage of 3.1 mm can be considered as a quite critical damage since the radius of the ball is 4 mm. The specifications of this type of bearing according to Fig. 1 are Nb = 9, Rb = 4 mm, Rc = 19.5 mm. The IM is energized from the power line (220 V/60 Hz). The Fig. 3 shows the damage used and the Fig. 4 shows the test base used to generate the database. A total of 100 signals are obtained by storing 614.4 cycles of the current signal, distributed in 65,536 samples; the signals of interest are obtained when the motor speed has reached the quasi stationary state, so from the 100 signals 2000 signals of 2560 samples each are obtained in this state. The current signal is acquired through 3 ACS758LCB-050B current sensors, conditioning the signal through a gain and DC-offset stage using the OP177; the ADS7841P converter digitizes the 3 phase currents and sends them by SPI protocol to a DILIGENT GENESIS 2 card that uses an FPGA of the Xilinx family" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001726_icems.2019.8921944-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001726_icems.2019.8921944-Figure3-1.png", + "caption": "Fig. 3. The positional relationship between stator magnetic potential and the d\u3001q axis", + "texts": [ + " ANALYSIS OF DEMAGNETIZATION PRINCIPLE The causes of demagnetization of permanent magnets are ambient temperature, high frequency vibration of permanent magnets and demagnetization potential of permanent magnets. This paper mainly analyzes the demagnetization of permanent magnets under demagnetizing magnetic potential; Under the same demagnetizing magnetic potential\uff0cThe difference between the demagnetization potential and the relative position of the permanent magnet has a great influence on the demagnetization size of the magnet. As shown in Fig.3. below. Demagnetization potential\uff1a When the stator magnetic potential coincides with the direct axis of the rotor, the demagnetization potential is the largest; when the stator magnetic potential is offset from the rotor magnetic potential, the stator magnetic potential decreases along the demagnetization direction component, and the demagnetization is weakened when the stator magnetic potential is intersected with the rotor. When the axial magnetic circuit is facing, the demagnetization potential of the magnetic field of the stator is zero" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002963_0021998320933664-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002963_0021998320933664-Figure1-1.png", + "caption": "Figure 1. Processing and assembling of the actuator. (a) Programming and surface treatment of the shapeshifting component. (b) Location and encapsulation of the reinforcement component. (c) Curing and sectioning of the actuator.", + "texts": [ + " At this point, to continue with the development of external stress-free cost-effective 2W-SMPs is necessary to satisfy the increasing demand for smart actuators and the control of their displacement. In this work, a new free-standing 2W- SMPs composite with great displacement capacity and better control of the actuation response is presented. A polyester urethane elastomer (PU) is used as an external stress-applying component, and a chemically treated surface crosslinked ethylene-vinyl acetate copolymer (cEVA) as the shape-shifting component, as shown in Figure 1 (a). The process of encapsulation of the programmed and treated shape-shifting component, as the second step process, enables the variation of the hardness and Young modulus of the matrix, enabling different shape-shifting responses. The actuation response depends on the capability of the cEVA to deform the actuator during the shrinkage (partial melting of the crystals), and the capability of the PU matrix provides the stress back to cEVA during the induced elongation (cooling below the crystallization temperature, Tc)", + " The reaction scheme and expected network structure of the crosslinked cEVA are shown in Scheme 1. Matrix stress-applying component: polyurethane elastomer. Stress-applying matrix was prepared by mixing a com- mercial urethane polyester, used as a prepolymer, and toluene diisocyanate as a curing agent (Smooth-on, Inc.), with 1:1 ratio. Samples were cured at 40 C for 12 hours to obtain a 30, 60 and 80 Shore A hardness polyurethane elastomer. Actuator assembling process. The assembly process of the actuator is schematized in Figure 1. First, the shapeshifting component is programmed and prepared for the encapsulation, cEVA-DCP dumbbell specimens according to the ISO 527\u20132 were cut from the cured sheets and programmed under 850 KPa at 95 C during 10min in a universal testing machine (UTM) with a customized thermal chamber. Then, the samples were cooled down to room temperature, and middle section of the programmed sample was sectioned to obtain bars of dimensions of 50mm 4mm 1mm. Next, the surface treatment was carried out by low plasma radiation during an exposition time of 300 s under a nitrogen atmosphere with negative polarity to improve Scheme 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001442_042023-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001442_042023-Figure1-1.png", + "caption": "Fig 1. Overall design drawing of foreign objects removal device for high- voltage transmission line The suspension part is divided into two parts: insulated hook and half-moon guide plate. The insulating hook is made of epoxy resin and made by 3D printing, which is convenient for factory processing. The front end of the insulated hook is equipped with a semi-moon-shaped suspension", + "texts": [], + "surrounding_texts": [ + "At present, the method of \u201cinsulated rope twisting\u201d is widely utilized in the removal of foreign objects in power lines. This method uses an insulating rope with a middle bolt spring-like wire hook to throw the wire, and the grounding worker pulls the insulating rope to the foreign object, and the twisting of the spring-shaped wire hook causes the foreign objects to be twisted with the rope. Then the foreign objects can be removed by pulling on the rope. The method of \u201cinsulated rope twisting\u201d has the following drawbacks. Firstly, in the case of solid foreign objects, this method often causes foreign objects and insulated rope hooks to entangle the wires, which make it difficult to remove foreign objects. Secondly, the wire hooks directly contact the lines. If the force is too vigorous, then the insulation performance of the insulated rope is decreased. Thirdly, for this method, it is difficult to control the force. Therefore, the use of this method may cause line oscillation and phase-to-phase short circuit, and also jeopardizes the safety of operator [3]. In this paper, a new foreign objects removal device used in high-voltage transmission line is proposed, which overcomes the shortcomings of traditional devices. The proposed foreign objects removal device is introduced in the next sections." + ] + }, + { + "image_filename": "designv11_80_0001739_ismsit.2019.8932756-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001739_ismsit.2019.8932756-Figure3-1.png", + "caption": "Fig. 3. Belt and Pulley Mechanism", + "texts": [], + "surrounding_texts": [ + "978-1-7281-3789-6/19/$31.00 \u00a92019 IEEE\nKeywords\u2014 automation, gardening, mobile application, timely monitoring\nI. INTRODUCTION\nIn manufacturing sector, to control machine tool a processed is used which is called CNC machining. The process includes the use of computers to control the machine tools. Tools can be any object which helps us in performing different task like extruder in 3D printer, blade in lathe machine, drill bit in milling machine. CNC stands for Computer Numerical Control. It look like an ordinary personal computer (PC), but it differs due to unique software and console to control the machine.\nTo control the speed, location, coordinates and feed rate of machine a specialized CNC machining language is used, which is known as G-code. With the help of CNC machining we can control the position and velocity with great precision. CNC machining is currently being used in manufacturing for plastic and metal parts.\nThere are lots of advantages of using CNC machines. It\u2019s hard to manufacture any part in manual machining. Manual machining consumes more time and energy and it does not have much precision. While using CNC machining a lot of man power and time is secured and also the parts manufactured with CNC machining has great precision.\nA CNC machine consist multiple individual motors, which helps its tool to move in respective specific direction. A CNC Machine consist of two or more axis of motion. Usually CNC machine which are being used are consist 3-axis of motion. If there is any additional rotational movement then it will be 4-axis.\nII. LITERATURE REVIEW\nCommercially, Agriculture has been reached a newly high level of automation up till now, mainly for growing\ncrops on broad land. Fine-grain satellite imagery is also available commercially for pesticides and fertilizer related applications, leading to the precision agriculture\u2019s novel paradigm, its basic goal is to save water and pesticides [1]. The interest is growing day by day in autonomous farming so we can move towards smaller robotic platforms that can work on individual basis and can done some manipulation in the field with the help of precise sensing [2].\nWith the increasing research in the field of computer vision [3] and mechatronics which is helping us to led to an autonomous solution for harvesting some specialty crops which includes Cherries [4], apples, tomatoes, cucumber, mushrooms, strawberries, melons and much other.\nAnother active research on which work is being done is automatic weed control. Grey-level vision is used to navigate in structured outdoor environment and uses color vision to differentiate between the required crops and weed which is needed to be removed. Beside their multiple applications in the field, we envision automated agriculture robot which is precise and able to work without any operator in any sort of environment like urban areas, house roofs or can easily work in harsh environment like outer space.\nAlso, a lot of work is done in this field like autonomous targeted spraying [5] to removes pests from crops which can destroy and can have an effect on crops production, design of optimum manipulator for autonomous de-leafing [6] process of cucumber plant to prevent it from fungal disease which is considered by its growers but is costly. In addition to all of this there are also some virtual experimentation frameworks which have been developed for agriculture robots, [7], [8] ForboMind is a customized software platform which was introduced to support and help field robot done agriculture task done with precision and to promote to reuse the robotics components.\nOn the other hand, Agriculture field robots contributes to improve soil health, yield and reliability of operation. Which are commonly equipped with multiple sensors and cameras for navigation, localization, mapping and path planning algorithm.\nAlso, farming industry make use of drones for surveillance and monitoring fields. These drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides.", + "III. SOCIO-ECONOMIC SIGNIFICANCE\nAt present farming industry make use of drones for surveillance and monitoring fields these drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides. Similar Machinery exists but is only household specific; however we are aiming to target research and development departments in pharmaceutical and food industry, where setting up small greenhouse is required. Further we would be using some tool head for multiple operations instead of relying on magnetism for tool selection. Cutting down cost is another objective by using alternative materials to steel.\nIV. DESIGN\nWhile developing the project we went through some major and minor issues which lead us to make multiple iterations in the design phase. In this section we will discuss these iterations.\n Iteration 1:\nThis is the initial idea and concept on which we started to work. Our initial idea was to have x-axis, y-axis, z-axis and one rotary axis for the axis of motion. We also thought about adding multi-head mounter and add moisture sensor, shower and seeder. Initially we want to use rack and pinion mechanism. Issues: Rack and pinion mechanism can increase the vibrations which can affect the structure.\nWe then start to work on to select a mechanism for x-axis. We start to work on the stepper motor and lead screw mechanism. Issues: Our project has an open base so lead screw can\u2019t be mounted at the bottom like every CNC milling machine and usage of two stepper motor for just one axis can effect or destroy the structure if something happens to one motor.\n Iteration 3: So we started to think more about it and started to work on another solution to move the x-axis with just one motor. We finally decide to move the whole axis with one motor by using belt and pulley mechanism as shown in figure below.\nIn this project we use t-slotted extruded bars for developing our structure. Y-axis consist of simple lead screw mechanism.\nZ-axis consist of simple lead screw mechanism for linear motion. On z-axis a servo motor is mounted for giving a rotary motion to the multi head. The final design after the iterations is shown below.", + " Vacuum Pump: When a signal is sent out from controller to the vacuum pump and vacuum pump turn on and suck air and move towards seed when it come near the seed, then seed float and got stuck on the nozzle until the vacuum is on. When it reaches to the desired location then the controller sends another signal and vacuum pump turns off and the seed drops on the location the grid.\n Water Pump: When a signal is sent out from the controller to the water pump then water pump turn on and sprinkle the water over the grid with the help of shower for some specific time and then another signal comes from the controller and turns it off. Piping and instrumentation diagram (P&ID) is shown below:\nCurrently there are lots of global challenges which we are facing like global warming, food needs, poverty, degradation of climate and much more. Sustainable development goals introduced by United Nation helps us address these problems and base on these problems find solution to make future better not just for humans but also for the creature that exist on globe. The goals which can be achieved by our project \u201cGarden Tech\u201d are as follows:\n Zero Hunger: As we all know that human population is increasing rapidly to accommodate them deforestation is taking place. It is effecting our climate and results in global warming due to which glaciers are melting. Flooding is taking place more often than ever before and leads to devastation of arable land. Due to these circumstances food growth is not increasing rapidly like human population. The anticipated population according to United Nations Department of Economic and Social Affairs (UN DESA) is shown below:" + ] + }, + { + "image_filename": "designv11_80_0003972_icma49215.2020.9233579-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003972_icma49215.2020.9233579-Figure2-1.png", + "caption": "Fig. 2 Design of a robotic wheel.", + "texts": [ + " Here, we estimate the maximum and minimum torque for the wheel locomotion. The total weight of the robot body is 4kgf and the radius of the pinion gear is 1.3 cm. The minimum required torque is estimated to be 5.2kgf-cm, and the servo motor is selected to meet the requirement. The specification of the servo motor (ROBOT SERVO RS301CRF3 FUTABA) is shown in Table 1. 1) Mechanical Design of the Wheel We design the robotic wheel by using a CAD software. The design and appearance of the entire body is shown in Fig. 2. The foundation is designed with a rail structure that allows the spokes to slide stably as shown in Fig. 3. The picture of the assembled robotic wheel based on the design drawing is shown in Fig. 1. The robotic parts are printed by a 3D printer (AGILISTA-3200, KEYENCE) and also fabricated by a modeling plotter (NC-5SK, Mimaki). The materials of the parts are acrylate resin and ABS. 2) Rack & Pinion Mechanism A pinion gear connected with a servo motor is combined with a rack gear with a spoke as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003131_j.optcom.2020.126271-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003131_j.optcom.2020.126271-Figure2-1.png", + "caption": "Fig. 2. Schematic cross-sectional structure and operating mechanism of the telescope objective: (a) The FOV without deflection. (b) The FOV with deflection. (c) The relationship between the interface deflection and the corresponding applied voltage electrodes.", + "texts": [ + " This device is a cubic chamber composed of four electrodes coated with hydrophobic insulator respectively. It is filled two immiscible liquids with equal densities but different refractive indices. One of the liquids is electrically conductive, the other is insulative. Especially, the conductive liquid contacts the substrate electrode. Different from the conventional liquid prism, a transparent partition sheet is placed between the L\u2013L interface. Therefore, a flat interface is formed in the cubic chamber. In the initial state, the FOV is the same as that of the conventional telescope, as shown in Fig. 2(a). When a voltage is applied between the conductive liquid and one of these sidewall electrodes, the contact angle between conductive liquid and sidewall is reduced due to electrowetting effect, as shown in Fig. 2(b). Then the L\u2013L interface is deflected, and the FOV is changed. In this way, we can get a larger FOV by deflecting the L\u2013L interfaces using different voltages as shown in Fig. 2(c). For our liquid prism, the position of the L\u2013L interface is changed due to electrowetting effect. According to the theory, there is a balance between the interfacial surface tensions (\ud835\udefe) near the three-phase conductive liquid (C), insulative liquid (I), and sidewall (S). In the initial state, these balances are related to the contact angle(\ud835\udf030) between the conductive liquid and sidewall by Young\u2019s equation: cos \ud835\udf030 = \ud835\udefe\ud835\udc46\ud835\udc3c \u2212 \ud835\udefe\ud835\udc46\ud835\udc36 \ud835\udefe\ud835\udc3c\ud835\udc36 (1) where \ud835\udefe\ud835\udc46\ud835\udc3c , \ud835\udefe\ud835\udc46\ud835\udc36 and \ud835\udefe\ud835\udc3c\ud835\udc36 are the interfacial tensions of the sidewall/ insulative liquid, sidewall/conductive liquid and insulative liquid/ conductive liquid, respectively. After the voltage is applied, according to Young\u2013Lippmann equation, the relationship of the contact angle \ud835\udf03 and the applied voltage U can be described as follows: cos \ud835\udf03 = cos \ud835\udf030 + \ud835\udf00 2\ud835\udefe\ud835\udc3c\ud835\udc36\ud835\udc51 \ud835\udc482 (2) where \ud835\udf00 is the dielectric constant of the hydrophobic insulator, d is the thickness of the hydrophobic insulator. Fig. 2(b) shows the optical pathway of beam passing through the liquid prism, which consists of two immiscible liquids with refractive indices denoted by \ud835\udc5b1 (insulative liquid) and \ud835\udc5b2 (conductive liquid). A mathematical relationship between the incident angle \ud835\udefc and the contact angle \ud835\udf03 between the conductive liquid and sidewall in liquid prism can be obtained as: \ud835\udefc = sin\u22121 \u23a1 \u23a2 \u23a2 \u23a3 \ud835\udc5b2 sin \ud835\udf03 cos \ud835\udf03 \u2212 \ud835\udc5b1 cos \ud835\udf03 \u221a 1 \u2212 ( \ud835\udc5b2 \ud835\udc5b1 cos \ud835\udf03 )2\u23a4 \u23a5 \u23a5 \u23a6 (3) To visually demonstrate the deflecting behavior of the proposed liquid prism, we first fabricate the device using transparent materials" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001184_j.mechmachtheory.2019.103606-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001184_j.mechmachtheory.2019.103606-Figure2-1.png", + "caption": "Fig. 2. Six combinations.", + "texts": [ + " 1 (c) shows a spatial quadrilateral chain quadrilateral with the identical beams (namely C p chain). Fig. 1 (d) shows a spatial upright triangle chain (namely C chain). Fig. 1 (e) shows a spatial inverted triangle chain (namely C chain). \u2207 Totally, 35 different parallel mechanisms can be obtained by randomly combining the limbs in Fig. 1 . There are 5 kinds of structures with three identical limbs, 20 with two same limbs, and 10 with three different limbs, as shown in Table 1 . Each combination in Table 1 represents a different parallel mechanism and Fig. 2 only lists six of them, where Fig. 2 (c) is a linear Delta parallel mechanism. 3. Kinematics analysis 3.1. Unified model For the sake of analysis, a schematic of the mechanism is provided in Fig. 3 , where a coordinate system O \u2212 xyz is located at the center of the base and another one P \u2212 x \u2032 y \u2032 z \u2032 is fixed at the center of the moving platform. r b and r a denote the radius of the base and the moving platform, respectively; L 1 and L 3 indicate a half of the length of the upper and lower beams for 4S mechanism, respectively; L 2 is the length of the fixed length struts", + " (4) Solving linear equations A k x k = \u2212b k , one can get the value x k . (5) Let x k +1 = x k + x k . (6) If \u2016 x k \u2016 \u2264 \u025b 2 , the output is x k +1 , and go to step (7) . Otherwise let k + 1 \u2192 k and x k +1 \u2192 x k ,and turn to step (2) . (7) End. The input can be solved by the inverse solution equation of the mechanisms after the corresponding x,y and \u03b1 are calculated for the given value of \u03b2 , \u03b3 and z . The dimensional parameters of the mechanisms are shown in Table 3 . Table 4 shows the results of the first configuration in Fig. 2 . Table 5 provides the results of the second configuration in Fig. 2 . Table 6 indicates the results of the fourth configuration in Fig. 2 . Table 7 demonstrates the results of the fifth configuration in Fig. 2 . It can be seen from Tables 4\u20137 that the each two of the six input parameters of the mechanisms are equal. This is in line with the actual situation since the two inputs are fixed on the same moving slider. By comparing the data in row 6, 7, and 8, it can be found that the three value x, y and \u03b1 do not change with the variation of z for the same value \u03b2 and \u03b3 . In addition, the three inputs change the same value. It can be concluded that the change of the value z is equivalent to shift the mechanism as a whole in the direction of the sliders and does not change the orientation of the moving platform" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002873_icpes47639.2019.9105570-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002873_icpes47639.2019.9105570-Figure4-1.png", + "caption": "Fig. 4: Stator models with integer-slot windings and different SPPs q.", + "texts": [ + " The stator model represents a 2D-model of an electrical machine with slots and will be used to identify the influence of the amplitude reduction of the harmonics. This model was presented in [2], where the focus was on the comparison of the losses between the rotor and the stator. This paper focuses on the influence of only one harmonic on the losses. In order to show the influence on the reduction of the amplitude of the single harmonic, three different stator models with the same pole pair number (p = 3) and different SPPs (q \u2208 {1, 2, 3}) are simulated. All three models are shown in Fig. 4. For the identification of the influence of the subharmonics on the losses, just fractional-slot concentrated windings with SPPs less than one (q < 1) are feasible. The stator is modelled with QS = 6 slots and wound with a double-layer concentrated winding (q = 0.5); the wound winding pattern is shown in the appendix in Fig. A-1. For a fair comparison the same stator is combined with rotors with different number of pole pairs. Thus depending on the rotor, the same geometry does or does not contain subharmonics", + " The absolute values of the stator, rotor and PM losses are presented in Table I as well. It can be seen, that the rotor losses induced by the sub-harmonics make up almost half of the total rotor losses. In addition to the analysis of the losses for every single harmonic, the impact of the amplitude reduction of the harmonics is considered in this section. The amplitude reduction is achieved at an integer-slot winding by increasing the SPP up to q = 3 (distributed windings) as well as reducing the coil-span by up to 2 slots (chorded windings). The used models were presented in Fig. 4. 1) Spectra: The simplest way to influence the amplitudes of the harmonics is to distribute the winding of one phase into several slots. The effect can be observed at the MMF spectrum of the considered fully-pitched integer-slot windings (q \u2208 {1, 2, 3}) in Fig. 9 . It can be seen, that increasing the SPP factor has a significant influence on the amplitudes of the harmonics, while the fundamental component has nearly the same level. Another method to reduce the amplitudes of the harmonics is chording", + " The sub-harmonics of the MMF distribution in the middle of the air gap have a substantial influence on the losses in the stator, rotor and PMs (Fig. 8). In order to better understand the sub-harmonic induced losses, an analysis of different models with different sub-harmonics is conducted. First, the sub-harmonics that occur by changing the number of pole pairs of the rotor model (Fig. 5) are considered. Secondly, the sub-harmonics of the single- and double-layer fractionalslot concentrated windings are analysed. 1) Changing Number of Pole Pairs: For the first study the stator model as shown in Fig. 4 with a number of slots QS = 6 is used. The corresponding winding pattern is shown Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 14,2020 at 15:36:02 UTC from IEEE Xplore. Restrictions apply. in Fig. A-1, and the rotor models in Fig. 5. The MMF spectra for all winding patterns are presented in Fig. 13. It can be seen that the MMF of the models with a pole number of p = 2 has no sub-harmonics. The winding scheme with p = 4 pole pairs has one and the model with p = 8 pole pairs has two sub-harmonics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000724_978-981-13-3305-7_190-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000724_978-981-13-3305-7_190-Figure1-1.png", + "caption": "Fig. 1. Overall configuration", + "texts": [ + " This section describes the mathematical model of the target distributed tilt-wing UAV, which consists of a nonlinear 6 degree-of-freedom rigid-body dynamic model, aerodynamic models of the VTOL and transition mode, including a propulsion model and a conventional fixed-wing aerodynamic model which considers the aerodynamic influence of slipstream produced by propellers. 2.1 The General Parameters and Configuration of the Target UAV As a preliminary technique confirmation of distributed propulsion tilt-wing aircraft, the target prototype UAV in this paper has four electric propulsion units along the wing, each of which includes a brushless rotor and a fixed propeller, and one tail electric propulsion unit whose thrust direction can deflect to a certain angle according to the attitude control strategy. The configuration is shown is Fig. 1. For future works, more electric propulsion units will be placed along the wing and tail to achieve the capability of heavy payload and redundancy propulsion control. The Rigid-Body Dynamic Model. A nonlinear 6 degree-of-freedom rigid-body dynamic model is established based on the Newton\u2019s law of motion, and established in the body axes system Oxbybzb and the Earth-fixed axes system Oxgygzg, that is, the North-East-Down (NED) system [7]. \u2022 Force equations for aircraft motion Fx \u00bc m\u00f0 _u rv\u00fe qw\u00de Fy \u00bc m\u00f0 _v\u00fe ru pw\u00de Fz \u00bc m\u00f0 _w qu\u00fe pv\u00de 8< : \u00f02:1\u00de \u2022 Moment equation L \u00bc Ix _p Ixy _q Ixz _r Ixzpq Iyzq2 \u00fe Izqr\u00fe Ixypr Iyqr\u00fe Iyzr2 M \u00bc Ixy _p\u00fe Iy _q Iyz _r\u00fe Ixpr Ixyqr Ixzr2 \u00fe Ixzp2 \u00fe Iyzpq Izpr N \u00bc Ixz _p Iyz _q\u00fe Iz _r Ixyp2 \u00fe Iypq Iyzpr Ixpq\u00fe Ixyq2 \u00fe Ixzqr 8< : \u00f02:2\u00de \u2022 Kinematic equation for rotational motion _/ _h _w 2 4 3 5 \u00bc 1 sin/tanh cos/tanh 0 cos/ sin/ 0 sin/sech cos/sech 2 4 3 5 p q r 2 4 3 5 \u00f02:3\u00de \u2022 Kinematic equation for translational motion _x _y _z 2 4 3 5 \u00bc coswcosh coswsinhsin/ sinwcos/ coswsinhcos/\u00fe sinwsin/ sinwcosh sinwsinhsin/\u00fe coswcos/ sinwsinhcos/ coswsin/ sinh coshsin/ coshcos/ 2 4 3 5 u v w 2 4 3 5 \u00f02:4\u00de Corresponding symbols and meanings are listed in Table 1: The Tilt-Wing Mechanism Dynamic Model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003143_jomms.2020.15.291-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003143_jomms.2020.15.291-Figure1-1.png", + "caption": "Figure 1. Schematic illustration of phase-evolution-based thermomechanical constitutive model reflecting process of phase transition between yellow rubbery phase and glassy phase in pewter.", + "texts": [ + " The proposed 3D constitutive model for SMPs with FEM implementation is applied to analyze two complex SMP structures and the outcomes are reported in Section 5. Finally, the concluding remarks are given in Section 6. Prior to the presentation of a 3D phase-evolution-based thermomechanical constitutive model formulation in Section 3, a 1D development that has been elaborated earlier by Li et al. [2017b] is briefly described in this section. 2.1. General description of the constitutive model. In this approach, SMPs are considered as a mixture of the yellow rubbery phase regions and the glassy phase regions in pewter color as shown in Figure 1. The volume fraction of each of the former is 1\u03b3 j r and the latter as 1\u03b3 j g , where the total sum is unity, i.e., 61\u03b3 j g +61\u03b3 j r = 1. The total strain of the adopted model including that of thermal effect is expressed as \u03b5total =61\u03b3 j r \u03b5 j rubbery+61\u03b3 j g \u03b5 j glassy+ \u03b5T (1) where \u03b5T is the thermal strain, \u03b5 j rubbery and \u03b5 j glassy are strains in any j-th rubbery and any j-th glassy phase regions, respectively. These two kinds of springs which represent two different phases (yellow for rubbery phase and pewter for glassy phase) are interchangeable governed by the variation of temperature", + " The strains in the glassy phase regions comprise both the mechanical strain \u03b5g-mechanics and the frozen strain \u03b5frozen. Hence, the total strain in Eq. (1) at the latter stage can be written as \u03b5total =61\u03b3 j r \u03b5 j r -mechanics+61\u03b3 j g ( \u03b5 j g-mechanics+ \u03b5 j frozen ) + \u03b5T (3) The stress in each phase region is described by Hooke\u2019s law as stipulated in Eqs. (4) and (5): \u03c3 j r = Er\u03b5 j r -mechanics, (4) \u03c3 j g = Eg\u03b5 j g-mechanics, (5) where Er and Eg are the Young\u2019s moduli of the rubbery and glassy phase regions, respectively. All spring elements shown in Figure 1 are in series and hence the stress in any phase region is the same and equals the total stress, namely, \u03c3total = \u03c3 j r = \u03c3 j g (6) As the mechanical strain of any j-th rubbery phase region remains unchanged, we have \u03b5 j r -mechanics = \u03b5r -mechanics. This is similarly true for the mechanical strain of any j-th glassy phase region and hence we also have \u03b5 j g-mechanics = \u03b5g-mechanics. Substituting these 2 identities into Eq. (3), the total strain becomes \u03b5total = (1\u2212 \u03b3 )\u03b5r -mechanics+ \u03b3 \u03b5g-mechanics+ \u03b5 f + \u03b5T (7) In Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001669_icems.2019.8921899-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001669_icems.2019.8921899-Figure6-1.png", + "caption": "Fig. 6. Sample of staotr bar in the involute segment for one stator slot.", + "texts": [ + " In order to study the influence of stator end structures on the leakage magnetic field for the end of stator bars, the end leakage magnetic field distribution is calculated by the threedimensional transient time-stepping finite element method when the relative permeability of the stator clamping fingers, the stator inner clamping plate, magnetic shield and the stator outer clamping plate of the physical model in Fig. 1 are 1 and the conductivity of that are 0. The centerline of the top layer bar in the involute segment is selected as sample line 1, the centerline of the bottom layer bar for the same stator slot in the involute segment is selected as sample line 2. The crosssection A is selected for the top layer bar and cross-section B is selected for the bottom layer bar in the involute segment. 40 sample points are selected for cross-section A and crosssection B, the sample point number is shown in Fig. 6. In sample line 1 and sample line 2, the fundamental RMS of the end leakage magnetic flux density when stator end structures are considered or not are compared in Fig. 7. The leakage flux density for the end of stator bars in different axial position is affected when considering stator end structures, which is larger because of the existence of the magnetic shield. The influence of stator end structures on the leakage magnetic flux density for the end region decreases along the direction far away from the stator iron core" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002922_icieam48468.2020.9112020-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002922_icieam48468.2020.9112020-Figure1-1.png", + "caption": "Fig. 1. Scheme of the crawling robot and the flight of stairs.", + "texts": [ + " A characteristic feature of the proposed robot is the possibility of combining caterpillar-, worm-, and snakelike movements, that expands its functionality and allows crawling onto various obstacles, including the ability to overcome a flight of stairs [11]. II. DESCRIPTION OF A CRAWLING ROBOT The flight of stairs, which is meant to be overcome by a crawling robot, is located in the absolute coordinate system so that the axis goes along the flight of stairs, and the axis \u2013 is perpendicular to it (Fig. 1). The flight consists of the -th number of steps, the length, width and height of each step are equal to lP, bP, and hP. Let us dwell on the study of the algorithm of the robot's crawling onto the first step of the flight. The radius vectors of the points A-H of the step equal to T( , ,0)A A A r T( , ,0)B A P A r T( , , )D A A P r T( , , )E A P A P r T( , , )F A A P P r T( , , )H A P A P P r The crawling robot considered in the present research is a three-link design, the links i=1-3 of which are telescopic (what is implemented using linear actuators, which are not shown in Fig. 1), their lengths are within the ranges min max[ , ]il l l , where lmin, lmax denote the minimum and maximum length values, whereby max min2l l . Each two links are interconnected by two active hinges 4 and 5. Hinges 4 provide the rotation of the links in the horizontal plane of the absolute coordinate system by the angles i, and the hinges 5 \u2013 in vertical planes of relative coordinate systems of the links by the angles i. The robot contacts the surface at four support points Oi, i=1-4, located at The work under consideration was carried out within the framework of the Presidential Grant -200", + " Authorized licensed use limited to: University of Exeter. Downloaded on June 26,2020 at 08:14:26 UTC from IEEE Xplore. Restrictions apply. the extreme points of the links. The peculiarity of the supports lies in the fact that they are equipped with two contact elements with different friction coefficients, at one of which (fmax) the support is fixed on the surface, and at the other (fmin) slides along it [12]: max min const, if , const, if . Oi Oi Oi f f r f f The change of the support surfaces is provided by linear drives, also not shown in Fig. 1. The height of the supports will be considered sufficiently small hi . In the following figures, the supports fixed to the surface are indicated by black triangles. The vector of generalized coordinates of the specified robot includes the coordinates of one of the support points, for example, the point 1, the link lengths li, as well as the rotation angles of the links in the horizontal i and vertical i planes T 1 1 1 1 2 3 1 2 3 1 2 3( , , , , , , , , , , , )O O O q The radius vectors of the control points can be written as follows" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001671_icems.2019.8921658-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001671_icems.2019.8921658-Figure1-1.png", + "caption": "Fig. 1 Cross section of BLDD machines. (a) Regular topology. (b) Proposed topology", + "texts": [ + " IV, a performance comparison of the regular and proposed BLDD machine is presented. Through magnetic circuit and FEA, it is proven that the proposed BLDD machine can reach higher torque density of PMSM part; At last conclusions will be drawn. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEEAuthorized licensed use limited to: University of Gothenburg. Downloaded on July 26,2020 at 23:59:17 UTC from IEEE Xplore. Restrictions apply. II. STRUCTURE AND OPERATION PRINCIPLE The cross section of the proposed BLDD machine is presented in Fig. 1 (b). The whole machine contains one stator and two coaxial. The stator has open slot, and two sets of windings are placed inside slots. Different from regular BLDD machine structure presented in Fig. 1(a), consequentpole Halbach magnets are set at the slot opening. The outer rotor is a flux modulator that consists of iron pieces, while the inner rotor has surface mounted magnets on the iron core. The stator windings are named as modulation winding and regular winding respectively, and their pole-pairs are set as: wm ro ri wr ro P P P P Z P (1) where Pro and Pri are the pole-pairs of the outer and inner rotor respectively, Z is the slot number, which equals to the pole-pair of consequent-pole Halbach magnets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.11-1.png", + "caption": "Figure 1.11. Inclinometer pendulum moving over the surface of the Earth", + "texts": [ + " This sensor gives the vertical direction. Traditionally, we use a pendulum (or a plumb line) for this. However, when we are moving, due to the accelerations the pendulum starts to oscillate and it can no longer be used to measure the vertical direction. Here, we are interested in designing a pendulum for which any horizontal acceleration does not lead to oscillations. Let us consider a pendulum with two masses m at each end, situated at a distance *1 and *2 from the axis of rotation of the rod (see Figure 1.11). The axis moves over the surface of the Earth. We assume that *1 and *2 are small in comparison to Earth\u2019s radius r. 1) Find the state equations of the system. 2) Let us assume that \" = \"\u0307 = 0. For which values of *1 and *2 does the pendulum remain vertical, for any horizontal movement of the pendulum? What values does *2 have to take if we let *1 = 1 m and we take r = 6 400 km for the Earth\u2019s radius? 3) Let us assume that, as a result of disturbances, the pendulum starts to oscillate. The period of these oscillations is called the Schuler period" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001384_s1068798x19100150-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001384_s1068798x19100150-Figure1-1.png", + "caption": "Fig. 1. Mockup of a small-diameter supporting roller.", + "texts": [ + " Thus, in each step, a linear problem is solved, but the stress\u2013strain state arising in the rubber component as a result of the deformation in the preceding steps is taken into account. The variation in the potential-energy increment at each deformation step is described by the expression [15] (11) In Eq. (11) where is the Treloar elastic potential; is a small parameter; and and are increments in the force vector and displacement vector, respectively, at the boundary of the region where the external forces are specified. As an example, consider the small-diameter supporting roller in Fig. 1. The annular rubber\u2013metal elements have an oval cross section (Fig. 2). The metal ring (thickness 1.5 mm) is connected to a rubber mass on vulcanization. We use IRP-1315 rubber. The stress\u2013strain state of the rubber components is determined in two stages. First, the stress\u2013strain state of the axial compression due to roller assembly is determined. In assembly, the rubber component is pressed in an external metal case and then axially compressed. In the second stage, the radial loading of the roller is considered", + " In these regions, the unit deformational energy is 567 and 560 kJ/m3, respectively. Most of the load in the given cross section is applied to the part of the rubber component between the isolines corresponding to = 67 kJ/m3. At the extreme point of rubber\u2013metal attachment, corresponding to the internal diameter (region 4), the concentration of the unit deformational energy is greatest (600 kJ/m3). Field tests of the small-diameter supporting roller with internal rubber shock-absorbing elements (Fig. 1) show that the life of the rubber components is 2000 h. The first fatigue cracks are observed in the regions where the unit deformational energy is concentrated. Subsequent operation is accompanied by peeling of the rubber from the metal ring in regions 2 and 4, temperature rise, and thermomechanical failure of the rubber components along the line connecting regions 1 and 4. The low unit deformational energy in regions \u0410 and B (Fig. 4) may be explained in that the design of the points where the metal rim is connected to the rubber element and the metal ring creates conditions for minimal deformation of the rubber" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002913_plans46316.2020.9109954-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002913_plans46316.2020.9109954-Figure1-1.png", + "caption": "Fig. 1. Diagram of CANS Lab ground based robot motor geometry", + "texts": [ + " The robots are constructed with a modular layered design allowing more sensors and capabilities to be added to the robots by simply adding new vertical layers while keeping the footprint of the robot the same. This modularity allows for flight-like hardware testing of numerous sensors such as vision sensors, unique communication electronics, range finding sensors, and inertial navigation sensors and eliminates the need to develop complex software models of these sensors. The base configuration of the robot is made up of three layers. The bottom layer houses the motor assemblies which consist of three stepper-motor assemblies evenly spaced 120 degrees around the robot, as shown in Fig. 1 Each motor is equipped with a \u201droller wheel\u201d assembly that allows for active motion in the direction of the wheel (denoted by v1, v2, and v3 in Fig. 1) as well as passive motion in the orthogonal direction of the wheels. This motor geometry coupled with the \u201droller wheels\u201d allow the robots to be omnidirectional, decoupling the translational movement of the robots from the rotational state and granting uniform movement in any direction. This omnidirectional property allows the robots to better simulate the motion of a space vehicle. The stepper motor assemblies are also equipped with encoders which allow for the measurement of each individual wheel\u2019s angular position and velocity state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001934_978-981-15-1124-0_17-Figure26-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001934_978-981-15-1124-0_17-Figure26-1.png", + "caption": "Fig. 26 Variations in total deformation of a femur bone", + "texts": [ + "4685e-004 6 8.9801e-004 7 449.56 8 523.51 9 1112.80 10 1382.50 The result from the modal analysis gives us the following 10 natural frequencies and mode shapes with fixed\u2013fixed boundary condition (Table 2 and Figs. 15, 16, 17, 18, 19, 20, 21, 22, 23, 24). Structural and Vibrational Analysis of Femur Bone Using FEA 541 Fig. 5 Mode Shape 1 Fig. 6 Mode Shape 2 The variations observed in the equivalent stress are shown in the Fig. 25. The variations in the total deformation of the Femur bone is shown in the Fig. 26. free\u2013free boundary condition gives us a frequency range from0Hz to 1382.50Hz and Fig. 7 Mode Shape 3 Fig. 8 Mode Shape 4 the analysis performed for the fixed\u2013fixed boundary condition gives us a frequency range of 1254.30 Hz to 8497.31 Hz. The results are in agreement with the literature taken for reference [1, 2]. The variations observed in the free\u2013free boundary condition can be accounted for by the fact that this is not a real case. Since the fixed\u2013fixed boundary condition gives us an accurate representation of the human body we can discuss the results of the fixed\u2013fixed analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001549_s11223-019-00110-8-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001549_s11223-019-00110-8-Figure1-1.png", + "caption": "Fig. 1. The structure and transmission principle diagram of the 2TDSRM toothed drive (a) and structural model of the single-tooth meshing pair (b).", + "texts": [ + " thermal strain and thermal stress distribution for the meshing pair model were obtained by thermal structural coupling analysis. Considering the thermal elastic deformation, we analyzed contact meshing stiffness, thermal meshing stiffness and comprehensive meshing stiffness and simulated the comprehensive meshing stiffness of different meshing positions. 1. The Structure and Transmission Principle. The 2TDSRM toothed drive consists of the movable tooth 1, ring gear 2, pin shafts 3, swing rod 4, separator 5, and wave generator 6, as shown in Fig. 1a; r0 is the base radius, c is the center distance, l is the swing rod length, and r1 is the movable teeth radius. When a wave generator rotates at an equal angular velocity, its profile face drives the movable teeth. At the same time, the movable teeth engage the ring gear which is stationary. Thus, the movable tooth rotates around the pin shaft under the constraint of ring gear profile. As a result, the movable toothed drive swings the swing rod and pushes the separator rotation connected with the output shaft. Speed transformation and power transfer of the drive takes place. The movable teeth form higher pairs with the wave generator and ring gear, respectively. As shown in Fig. 1b, one motion cycle of the drive is analyzed as follows: from the beginning contact point W, the movable tooth moves along the surface e of the ring gear from the top to root driven by the lift curve of the wave generator. At the ending point V , the working stroke of the meshing is finished. From point V to U , the movable tooth moves along the ring gear surface f driven by the separator, this process is called return stroke. 2.1. Analysis of the Heat Flow Density of the Meshing Pair. The heat flow density of the meshing pair of the ring gear and movable tooth depends on the relative sliding velocity, meshing contact stress, heat flux distribution, friction and thermal energy conversion coefficients" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001252_systol.2019.8864796-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001252_systol.2019.8864796-Figure1-1.png", + "caption": "Fig. 1: The earth frame E and the body frame B", + "texts": [ + " Section III analyzes the disturbance torque generated in the transient phase, if the standard degraded controller is used. In Section IV, the design of the modified degraded controller is introduced. In Section V, the results of simulation are presented. Finally, some concluding remarks are given in Section VI. I I . DY NA M I C M O D E L For the design of a controller, it is important to understand the dynamics of the system to be controlled. The mathematical model derived in [10] is recapitulated in this section. The typical configuration of hexacopter shown in Fig. 1 has six rotors, denoted by ri, i = 1, ...,6, respectively, which are perpendicular to the corresponding arm and generate the propeller forces Fi,z, i = 1, ...,6 and the reaction torques \u03c4i,z, i = 1, ...,6 [11]. An earth frame E and a body frame B are needed to describe the motion of hexacopter. The position and the orientation of hexacopter in the earth frame are defined, respectively, as \u03beE = [x y z]T and \u03b7E = [\u03c6 \u03b8 \u03c8]T , where 978-1-7281-0380-8/19/$31.00 \u00a92019 IEEE 171 \u03c6 , \u03b8 , \u03c8 represent the Euler angles, which are called roll angle, pitch angle and yaw angle, respectively", + ",6 can be calculated by (4), obviously the propellers have the same rotation speed. Due to small roll and pitch rate (i.e. \u03c6\u0307 \u2248 0, \u03c6\u0308 = 1 Ix u2, \u03b8\u0308 = 1 Iy u3, \u03c8\u0308 = 1 Iz u4. (8) During Stage 1, it can be assumed that the speed of rotors r2, r3, r4, r5, r6 are identical, the reaction torques of r2, r3, r5, r6 compensate each other. The thrust force F4 of rotor r4 generates a torque \u03c42,1, which causes a tilted angle along YE -axis \u03b81 = 1 Iy \u222b\u222b t1 0 \u03c42,1dtdt and a reaction torque \u03c43,1 = kM lkT \u03c42,1. The first part of the disturbance torque is \u03c4dis,1 = \u03c43,1. According to Fig. 1 the direction of reaction torque \u03c43,1 is clockwise. During Stage 2, the FDI system has already detected the fault and isolated the faulty actuator in rotor r1, simultaneously the corresponding rotor r4 is stopped. The standard degraded controller is applied to control the rest rotors r2, r3, r5, r6 to stabilize the position of hexacopter. To compensate the angle \u03b81, the speed of rotors r2 and r6 are increased and the speed of rotors r3 and r5 are decreased. In this case, the opposite torque \u03c42,2 and the reaction torque \u03c43,2 can be described by \u03c42,2 = 1 2 lkT (\u03c9 2 3 +\u03c92 5 \u2212\u03c92 2 \u2212\u03c92 6 ) and \u03c43,2 =\u2212kM(\u03c92 3 +\u03c92 5\u2212\u03c92 2\u2212\u03c92 6 )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000962_icma.2019.8816321-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000962_icma.2019.8816321-Figure1-1.png", + "caption": "Fig. 1. Inertial frame and body-fixed frame definition of UVMS", + "texts": [ + " Finally, the excellent robustness, rapidity and smoothness of the controller are verified by MATLAB simulation experiments. The paper is organized as follows. Section 2 gives a brief description of the UVMS. Section 3 presents the main results of the paper. Afterwards, the simulation of the UVMS controller is conducted and corresponding results are given and analyzed in Section 4. Finally, our contribution is summarized in Section 5. The simplified model of a 10-DOF UVMS with a 6-DOF AUV and a 4-DOF manipulator, as shown in Fig. 1, can be expressed (detailed derivation can be found in ): ( ) ( , ) ( , ) ( , )+I B dM q C q D q G q R\u03b6 \u03b6 \u03b6 \u03b6 \u03b6 \u03c4 \u03c4+ + + = (1) kJ\u03ba \u03b6= (2) Where 1 2 3 4[ , , , , , , , , , ]Tx y z\u03b7 \u03c6 \u03b8 \u03c8 \u03b8 \u03b8 \u03b8 \u03b8= stands for the position and orientation vector of the system in the inertial frame and joint position, 1 2 3 4[ , , , , , , , , , ]Tu v w p q r\u03b6 \u03b8 \u03b8 \u03b8 \u03b8= is the linear and angular velocity vector of the system in the body-fixed frame and joint velocity. ( )M q , ( , )C q \u03b6 , ( , )D q \u03b6 stand for inertial matrix, Coriolis and centripetal forces, hydrodynamic damping term of the system, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003687_978-3-030-58799-4_19-Figure17-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003687_978-3-030-58799-4_19-Figure17-1.png", + "caption": "Fig. 17. Courant number mean - arc-shaped case", + "texts": [ + " Apparently, the horizontal tailplane provides a negative lift when the elevator is not deflected, due to its initial negative incidence with respect to the aircraft centerline (ih < 0) and the downwash effect from the wing. Deflecting the elevator by \u221210\u25e6 increases the absolute value of the negative lift. On the other hand, deflecting the elevator by +10\u25e6 provides a positive lift, that is in absolute value lower than the previous case. The drag polar obtained for \u03b4e = \u221210\u25e6 is shown on the right side of Fig. 16. It is worth noting that the negative deflection of the elevator leads to a relevant increment of 78 drag counts for CD0. By looking at the drag breakdown, which is reported in Fig. 17, it is interesting to note CFD Prediction of Aircraft Control Surfaces Aerodynamics 85 Y cC l / c re f -1 -0.5 0 0.5 1 0 0.2 0.4 0.6 0.8 no deflection 10 deg deflection Fig. 11. Ailerons: spanwise wing loading at zero AoA with, and without ailerons deflection. CL C r 0.2 0.4 0.6 0.8 1 1.2 1.4 -0.045 -0.04 -0.035 -0.03 CFD VLM Fig. 12. Ailerons: rolling moment versus lift coefficient. that deflecting the elevator by \u221210\u25e6 (up) leads to higher drag and influences the vertical tailplane as well. This is due to the fact that the elevator leads to higher (negative) load in this case", + " NPB Job Requirements Job CPU Memory (MB) Runtime (sec) BT 2 1280 180 FT 2 5132 240 MG 4 26624 420 SP 6 5132 680 We run 110 jobs for the experiments: 34 BT, 34 FT, 25 MG and 17 SP jobs. The number of jobs with small requisitions was large, so the scheduling algorithms were able to perform the advance of a large number of jobs. Like in previous experiments, the cluster was used in a dedicated way and each algorithm was executed 30 times. For the BT, FT and MG jobs the standard input data was used. For the SP jobs, the number of iterations was changed to 160 and the problem size to 256 \u00d7 256 \u00d7 256. Figure 17 illustrates the makespan when using the four algorithms in this case. It is possible to observe that the EASY-backfilling algorithm presented improvement in execution time by 10% when compared to the SJF algorithm. It seems that EASY-backfilling algorithm respected the execution queue and executed the largest jobs in its intended time, unlike the SJF algorithm, which advanced smaller jobs and left larger jobs to the end. As a result, larger jobs end up delaying execution. Also, in this scenario, the FIFO achieved better performance when compared to the SJF, precisely by running the larger jobs in their time and not delaying them at the beginning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002121_3373724.3373741-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002121_3373724.3373741-Figure2-1.png", + "caption": "Figure 2. Rotating Vector", + "texts": [], + "surrounding_texts": [ + "In the earth frame, translational motion is easy and obvious. \ud835\udc39 mV 12 V \ud835\udc4b\ud835\udc4c\ud835\udc4d 13 In which, P is the momentum, and F is the net force implemented upon object. They are both expressed in the earth frame. Force vector related to earth equals to the force vector in body frame with left multiply \ud835\udc45 . \ud835\udc39 \ud835\udc45 \ud835\udc39 \ud835\udc45 00\ud835\udc47 00\ud835\udc5a\ud835\udc54 \ud835\udc47 \ud835\udc36 \ud835\udc36 \ud835\udc46 \ud835\udc46 \ud835\udc46\ud835\udc47 \ud835\udc36 \ud835\udc46 \ud835\udc46 \ud835\udc36 \ud835\udc46\ud835\udc47\ud835\udc36 \ud835\udc36 \ud835\udc5a\ud835\udc54 14 \ud835\udc4b\ud835\udc4c\ud835\udc4d \u239d\u239c \u239b \ud835\udc54 \u23a0\u239f \u239e 15 For rotational motion, it is important to figure out the mathematical description of a rotating vector. Assume an arbitrary X vector in a 3D space. Let this vector rotate about an arbitrary axis, we define this rotation vector as \u03c9. Using right-hand rule, we can get a delta X created by rotating. \ud835\udc4b is a vector directs from end of X to end of X(t+1). Then we can have \ud835\udc4b \u03c9 X# 16 Assume this vector as X axis in a X,Y,Z frame, and this frame rotates by a velocity \u03c9. \ud835\udc4b \u03c9 X \ud835\udc4c \u03c9 Y 17 \ud835\udc4d \u03c9 Z \u03c9 \u03c9\u03c9\u03c9 18 To make it simple and matrix form, \u03c9 is invented. With using \u03c9, the product can be eliminated. \u03c9 \ud835\udc36\ud835\udc5f\ud835\udc5c\ud835\udc60\ud835\udc60 \ud835\udc5d\ud835\udc5f\ud835\udc5c\ud835\udc51\ud835\udc50\ud835\udc62\ud835\udc61 \ud835\udc42\ud835\udc5d\ud835\udc52\ud835\udc5f\ud835\udc4e\ud835\udc61\ud835\udc5c\ud835\udc5f 0 \u03c9 \u03c9\u03c9 0 \u03c9\u03c9 \u03c9 0 19 \ud835\udc4b \u03c9X \ud835\udc4c \u03c9Y 20 \ud835\udc4d \u03c9Z {\ud835\udc4b, \ud835\udc4c, \ud835\udc4d} is also defined as \ud835\udc45, therefore we get transform matrix between R and R . [7] R \u03c9R 21 Torque acting upon copter is M, which equals to the time derivative of H, the angular momentum. I state the inertia matrix of this quadcopter. \ud835\udc3c \ud835\udc3c 0 00 \ud835\udc3c 00 0 \ud835\udc3c 22 H I \u2219 \u03a9# 23 Remember torque and thrust implemented upon frame are led by propellers which are all in body frame. So, we need to involve M rather than M in the equations. M I\u03a9 \u03a9 I\u03a9 24 \u03a9 \ud835\udc3c M \u03a9 \ud835\udc3c\u03a9 25 \u03a9 \ud835\udc5d\ud835\udc5e\ud835\udc5f \u239d\u239c\u239c \u239b \u23a0\u239f\u239f \u239e 26 Add translational dynamics and rotational dynamics equations together, we get the total dynamics equation as follow. [8] \u239d\u239c \u239c\u239c\u239b \ud835\udc4b\ud835\udc4c\ud835\udc4d\ud835\udc5d\ud835\udc5e\ud835\udc5f \u23a0\u239f \u239f\u239f\u239e \u239d\u239c \u239c\u239c\u239c \u239c\u239c\u239c \u239b \ud835\udc54 \u23a0\u239f \u239f\u239f\u239f \u239f\u239f\u239f \u239e 27" + ] + }, + { + "image_filename": "designv11_80_0000339_17686733.2019.1592392-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000339_17686733.2019.1592392-Figure2-1.png", + "caption": "Figure 2. (a) CAD model of test structure (cross-sectioned) (b) Schematic of the powder nozzle with mounted thermal imaging cameras in coaxial configuration on the top and the side view configuration.", + "texts": [ + " In comparison to well-established powder bed systems such as SLM systems, the laser cladding systems can build up larger volumes [5] and the wastage of the cost-intensive metal powder is significantly lower. Test structure and cladding parameters The test structure was designed to demonstrate the capabilities of the developed system. Therefore, the layers of the workpiece should vary in shape, and the build-up of the structure should contain typical challenges for the laser cladding process and the system. The test structure we chose for this work consists of a hidden pyramid-shaped cavity within a 20 x 20 x 20 mm3 cube. A cross section of the structure as CAD model is illustrated in Figure 2(a). The challenges for this structure are the tilted pyramid walls and the closing of cavities such as the pyramid tip. As material, a 316L stainless steel powder was used. In contrast to metal powders such as nickel-chromium-based superalloys, manufactured structures made of the 316L steel do not show cracks after assembly [10]. The process for one test structure was done in approximately 135 min. Further process and test structure details are given in Table 1. Figure 1. Schematic of a laser cladding system", + " Melt pool temperature measurement and nozzle movement logging The measurement of the melt pool temperature was performed using a PI 05M thermal imaging camera developed by the company Optris GmbH. This is a camera specially designed for temperature measurements of molten metals. Due to its spectral range of (500\u2013540) nm, it can handle the low emissivity of reflective metal surfaces and shows a high-temperature sensitivity at (900 to 2000) \u00b0C and an image format of 382 \u00d7 288 pixel at a frame rate of 80 Hz. Figure 2(b) shows the two tested arrangements of the camera at the laser cladding system. In the first assembly, the camera wasmounted at an angle of 45\u00b0 and at a distance of 15 cm from the melt pool. This has the advantage of a wide view across the manufactured component and a more detailed image of the melt pool due to the smaller distance to it. In the second assembly, the camera was coaxially mounted to the optical path of the laser beam on the top in a distance of 30 cm to the melt pool surface. The laser focus lens in the optical path inside the powder nozzle causes a slight magnification of the camera image and other optical components such as themirror/beam guide and the lens itself are causing transmission losses which need to be considered" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003479_nano47656.2020.9183696-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003479_nano47656.2020.9183696-Figure5-1.png", + "caption": "Fig. 5. Schematics of nanorobotic fabrication and control. (a) Illustration of a Janus particle; (b) SEM image of the Janus particle; (c) Schematic of magnetic control.", + "texts": [ + " The experimental platform contains a coil system, three power supplies, a National Instruments data acquisition controller, an inverted microscope, a camera, and a host computer. The power supplies are programmable and can generate sinusoidal outputs to the coils to create a rotating magnetic field. The camera provides visual feedback and records the location information of Janus particles. The computer is used as a control interface for the camera and the DAQ controller [9, 10]. Janus particles are fabricated by half-coating polystyrene (PS) particles, which are 10 \u03bcm in diameter, with a 100 nm thick layer of nickel (Ni) using electron beam evaporation as shown in Fig. 5(a). The surface morphology of the Ni-PS particles is confirmed with scanning electron microscopy (SEM) as shown in Fig. 5(b). For translational motion control of the nanorobots, we apply rotating magnetic fields generated by a 3D Helmholtz coil system. The xy motion of the nanorobot is controlled by manipulating the strength (mT), orientation (rad), and frequency (Hz) of the rotating magnetic field. The generated magnetic field can be expressed as ?\u20d7\u20d7? = [ \ud835\udc35\ud835\udc5fsin\ud835\udf03cos\ud835\udf14\ud835\udc61 \ud835\udc35\ud835\udc5fcos\ud835\udf03cos\ud835\udf14\ud835\udc61 \ud835\udc35\ud835\udc5fsin\ud835\udf14\ud835\udc61 ], (9) where \ud835\udc35\ud835\udc5f , \ud835\udf03, \ud835\udf14, and t represent the magnetic field strength, direction of motion, rotational frequency of the field, and time, respectively (Fig. 5(c)). In the in vitro experiment, three nanorobot swarms were employed sequentially in the injection position to target the three tumors one by one. Specifically, after a tumor is detected by a nanorobot swarm, the BGF landscape of the search space will be modified according to the BGF information gathered during the first TST period. Then, another nanorobot swarm will be employed to detect the second tumor according to the modified BGF landscape. Following this way, all the three tumors in the search space 364 Authorized licensed use limited to: Carleton University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000243_s12666-019-01662-8-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000243_s12666-019-01662-8-Figure4-1.png", + "caption": "Fig. 4 Schematic of graphite die and punch assembly used for temperature gradient sintering", + "texts": [ + " It is to be noted that the *10%, 50%, and 90% of the particles in the powder sample are smaller than the particle diameter denoted by D10, D50, and D90, respectively set pressure is attained by the time the graphite die temperature rises to 900 C. The graphite die temperature (3 mm away from the inner wall) was measured at the center (i.e., close to layers L3 and L4) using a K-type thermocouple (1.5 mm diameter). Figure 3b shows a schematic of the sintering cycle. The sintering conditions were arrived at after extensive trials. In the second round, SPS experiments were conducted using a temperature gradient die, as shown in Fig. 4, to achieve satisfactory densification in all the layers without causing excessive heating and undesirable intermetallic formation. The graphite die was designed such that, for a given power input, the temperatures attained in the powder sample were the highest in layer L6 (100 vol.% SS) and lowest in layer L1 (100 vol.% Ti). Similar die designs were used for better effect by Tokita [22], Jin et al. [23], and Hong et al. [24] for producing ZrO2/stainless steel, TiB2/ AlN/Cu, and ZrB2\u2013SiC/ZrO2(3Y) functionally graded materials, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001218_oceanse.2019.8867134-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001218_oceanse.2019.8867134-Figure2-1.png", + "caption": "Fig. 2: The diagram shows the development process of the fin drives. On an oscillating drive system (a), the fin shapes and materials (b) determined in a screening process are first examined for thrust development in the water channel. Favourable combinations of characteristics are then integrated into the ROV in the form of fin-like drive elements (c).", + "texts": [ + " The results from the numerical flow provide additional information for the detailed optimization of tear-off edges and partial structures (Fig. 1). In further project steps, alternative propulsion systems based on fish fins and biological movement patterns were investigated. Analogous to the representations for the hull optimization, biomimetic selection procedures were carried out to determine the optimal constellation of fin geometries and stiffnesses at changing frequencies and amplitudes (Fig. 2). Based on the brushless outrunner motors (DST 700, Turnigy, rpm/v = 700) of the OpenROV, a 3D printable mechanic was developed to generate an oscillating motion. We used a worm gear (reduction ratio 25:1) to get oscilating frequencies between 1-5 Hz at an amplitude of 68\u00b0. Fish inspired fins were 3D printed in ABS with different aspect ratios (AR = 0.4-5.6, A=5000 mm\u00b2, thickness 0.5 mm), fin geometries and stiffness gradients. The thrust measurements were carried out in a water channel (vmax 1 m/s, measuring section: 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001429_j.mechmachtheory.2019.103674-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001429_j.mechmachtheory.2019.103674-Figure15-1.png", + "caption": "Fig. 15. Application of the optimisation-based haptic synthesis and incremental changes of the virtual prototype.", + "texts": [ + " Such a translation in the horizontal plane does not affect the haptic properties since the user can change his position equivalently but it allows to continue the haptic synthesis. It can be activated through a double clicking metaphor, applying two consecutive force impulses at the handle in vertical direction. The new position is then determined through an optimisation, maximizing the clearance between the coupler curve and the workspace boundaries, and updated in the haptic simulation. The application of the haptic synthesis is shown in Fig. 15 . Within 21 s a user applies some normal forces to change the virtual prototype to his needs. Incrementally the initial mechanism is adapted during interaction and the user can directly feel the kinematic and kinetostatic properties of the mechanism. Fig. 15 a shows the initial and the final mechanism as well as some coupler curves of mechanisms in between. In Fig. 15 b the time behaviour of some related link lengths are shown. This paper introduces the haptic synthesis of mechanisms for the first time and discusses all required components. The motivation for the haptic synthesis, i.e. designing high-quality hand-actuated mechanisms using virtual prototyping, and the integration into the mechanism design process is shown in Section 1 . The modelling and control of the HFS is described in Section 3 . An exemplary virtual prototype, i.e. a hand-actuated four-bar with spring support and limit positions, is introduced in Section 4 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001903_j.jmmm.2019.166372-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001903_j.jmmm.2019.166372-Figure3-1.png", + "caption": "Fig. 3. A sample with windings. The maximum flux density measured by the search coil was controlled at 1.4 T.", + "texts": [ + " The die used for producing these samples is the same as that for ring cores shown in [6], which restricts the structure of the samples. Wire cutting was applied to cut out single sheets and laminated cores to minimize the magnetic degradation in the cutting process. The rolling direction is horizontal in this figure. The width of these cores is 6 mm, the same as that of ring cores shown in [6]. The length and width of each interlock, a and b respectively, are 4 mm and 1 mm. The overall height of the stack T is approximately 5 mm. The height of the protrusion shown in Fig. 3(c) is 0.35 mm. The photo of a sample with windings is shown in Fig. 3. Fig. 4 shows Bmax, Hmax, and W distributions for a single sheet without interlocking. Bmax and Hmax are defined by the following equations. = +B B Bmax{ }x ymax 2 2 (4) = +H H Hmax{ }x ymax 2 2 (5) The frequency of excitation was 50 Hz. The maximum flux density measured by the search coil shown in Fig. 3 was controlled at 1.4 T. This measurement has 86 and 3 measurement points for longitudinal and transverse directions, respectively. As shown in Fig. 5(a), the distances between the adjacent measurement points are approximately 1 mm and 0.75 mm for longitudinal and transverse directions, respectively. In Fig. 4, the measurement results for each points are shown by colored rectangles. The rectangles are located at the center of the measurement points. Although the size of the sensor restricts the measurable area, local magnetic characteristics near the cut edge (or the interlocking shown below) can be deduced by comparing neighboring points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001788_1350650119895192-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001788_1350650119895192-Figure2-1.png", + "caption": "Figure 2. The experimental test rig.", + "texts": [ + " However, if the journal bearing is allowed to adjust to the proper attitude angle according to the tilt of the shaft, then the oil film will be evenly distributed and the bearing will be working under better condition. As a result, the journal bearing will become stable and get a longer life. The bearing of this kind with a pre-tilt angle is defined as pre-tilted journal bearing. A test rig has been designed and built to validate that the journal bearing with a pre-tilted angle will have better performance. The whole test rig can be seen in Figure 2. A rotor with two disks is supported by two bearings. The bearings are fixed in the jigs. The design of the jigs, however, as the figure shows, is different between the jig #1 and jig #2. Thus, the experiments are conducted with the different two jigs for comparison. The jig #1 is designed to be fixed in the bearing support and could not move during the experiment. While the jig #2 is fixed in the inner ring, of which the outer surface is convex spherical surface. The outer ring is fixed in the bearing support, and the inner surface of the outer ring is designed to be concave spherical surface, thus forming the mating surface couples with the inner ring" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000090_052079-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000090_052079-Figure13-1.png", + "caption": "Figure 13. The use of temporary support structures and assembly sequence of reciprocal grillage.", + "texts": [ + " The beams according to figure 11 make it possible to construct grillages with a free-form configuration based on a design grid with module m (figure 12). WMCAUS 2018 IOP Conf. Series: Materials Science and Engineering 471 (2019) 052079 IOP Publishing doi:10.1088/1757-899X/471/5/052079 WMCAUS 2018 IOP Conf. Series: Materials Science and Engineering 471 (2019) 052079 IOP Publishing doi:10.1088/1757-899X/471/5/052079 An issue requiring a separate analysis is the method of making the grillage from short beams, including the order of assembly of individual elements and the solution of supporting the structure for the time of assembly (figure 13). Since none of the beams is supported solely on the walls or columns, each requires a temporary support. The assembly should be started from the beam B1, which, in one end, rests on the target support, which allows the use of only one temporary support supporting the other end of the beam, and also allows control of the beam position in a horizontal plane of an WMCAUS 2018 IOP Conf. Series: Materials Science and Engineering 471 (2019) 052079 IOP Publishing doi:10.1088/1757-899X/471/5/052079 established ordinate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002696_s40430-020-02368-5-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002696_s40430-020-02368-5-Figure3-1.png", + "caption": "Fig. 3 Cutting gearbox housing", + "texts": [ + " Therefore, the reliability of coal mining machine is one of the most concerned problems of coal mining companies. However, various unexpected problems have been observed during the cutting operation, and CPGH is one of the weak parts [1]. Recent problems [2, 3] involve the fracture failure of housing during operation in section A\u2013A and B\u2013B (Fig.\u00a02), resulting in reduction in coal production. CPGHs are different from traditional housings (automotive, aerospace, industrial, wind turbine gearboxes, et\u00a0al.) which are designed to be like cantilever beams for the objective of coal cutting as shown in Fig.\u00a03. A CPGH consists of motor shell, middle part of gearbox housing and planetary gear system shell (Fig.\u00a02). The excitations of CPGHs involve not only high-frequency internal vibrations caused by gear meshing, but also low-frequency heavy external vibrations caused by coal cutting. Stress concentration and excessive deformation can be easily caused by the cantilever beam shape of CPGHs. Previous Technical Editor: Wallace Moreira Bessa, D.Sc. * Yimin Zhang zhangyimin@syuct.edu.cn 1 School of\u00a0Mechanical Engineering and\u00a0Automation, Northeastern University, Shenyang\u00a0110819, China 2 Equipment Reliability Institute, Shenyang University of\u00a0Chemical Technology, Shenyang\u00a0110142, China 3 Shenyang Engine Design Institute of\u00a0China Aero Engine, Shenyang\u00a0110000, China 1 3 Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:273 1 3 Page 3 of 15 273 papers [4\u20137] indicated that fatigue failure is due to the joint action of local resonance and stress concentration and the failure region is often the part of stress concentration and large strain in local resonance", + " This paper presents a combination study of force analysis, finite element analysis, and field tests to investigate the main causes of the fatigue failure of coal mining machine (CPGH). Force analysis is proposed to obtain the regions of stress concentration under external excitation, and stress values are calculated by finite element analysis. Field testing and modal analysis are employed to get the resonance frequencies and main vibration patterns under internal excitation. Further investigation of the influence of key design parameters on stress and natural frequencies of gearbox is proposed for avoiding the resonance frequencies and reducing stress concentration. Figure\u00a03 shows the overall mechanical model of gearbox housing with external excitation caused by coal cutting. The regions of stress concentration and stress state of them are obtained by this model. The equations for gearbox housing are: where Mp is external torque acting on gearbox housing which is caused by meshing force between planet gears and ring gear, as shown in Fig.\u00a04. This torque was not considered by the previous papers [17\u201320, 23]. The principle of Mp is dynamic meshing force ( Fm ) which can be decomposed into radical ( Fy ) and circumferential directions ( Fx ) will be generated with the process of meshing between planet gears and ring gear", + " CPGH is a integral casting housing, and the design aim is to satisfy strength and stiffness conditions under both internal and external excitations. The design parameters D, B and H provide space of transmission system, and they are usually regarded as constant due to the limited space of coal mining working face. Therefore, wall thickness of mining goaf side (b1), mining face side (b2), upper and lower wall thickness (h), planetary gear system shell wall thickness (p), rib thickness (l) are regarded as design variables, and all of them are shown in Fig.\u00a03. Here, the range of b1, b2, h, k is 40\u201370\u00a0mm, and p is 65\u201395\u00a0mm (based on design drawings). Finite element analysis was made at the intervals of 2\u00a0mm, and the influence of design parameters on natural frequencies and stress amplitude was obtained. 1 3 Figure\u00a018 shows the influence of design parameters on stress concentration. It can be observed that the influence of design parameters is: b2 > h > l > b1, and the key parameter is b2. The stress amplitude value decreases quickly with b2 increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003616_j.matpr.2020.08.536-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003616_j.matpr.2020.08.536-Figure8-1.png", + "caption": "Fig. 8. Total Deformation.", + "texts": [ + "45 mm & 3D Number of Elements 27,983 Solver Sparse Direct Number of Nodes 42,662 develop a system before it is being manufactured. We can resolve the errors which are made during the design phase which avoids defects in the components (Fig. 7). The component is studied before it gets manufactured which reduced the occurrence of errors made. Flow simulation can also be performed and can get the results even there are flow simulation features which can be selected prior to the simulation. It also shows the results displayed on the screen (Fig. 8). An effective analysis to our Gear arrangement and metal casing to withstand all sorts of it takes, we have used Ansys workbench 16.0 we have done the static structural analysis by fixing the spindle and given the dynamic load at track arm with the Triangular mesh. The results obtained are within the limits of factor of safety. Worm and Wheel Analysis: - The various tests done (Table 5) for the Worm and Wheel Gear: - Total Deformation Directional Deformation Shear Elastic Strain Von-mess Stress Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003112_j.mechmachtheory.2020.103903-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003112_j.mechmachtheory.2020.103903-Figure2-1.png", + "caption": "Fig. 2. Hybrid serial-parallel 5-DOF 2-coupled-Cartesian-manipulator with non-parallel-revolute-axes.", + "texts": [ + " Stuart [53] reports the manipulators in Figs. 2 , 8 . The manipulators in Figs. 3\u20136 , 9 , 10 are novel. Please cite this article as: P. Wiktor, Coupled Cartesian manipulators, Mechanism and Machine Theory, https://doi.org/10. 1016/j.mechmachtheory.2020.103903 8 P. Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx T Please cite this article as: P. Wiktor, Coupled Cartesian manipulators, Mechanism and Machine Theory, https://doi.org/10. 1016/j.mechmachtheory.2020.103903 P. Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx 9 Fig. 2 hybrid serial-parallel 5-DOF 2-coupled-Cartesian-manipulator with non-parallel-revolute-axes. A common-link L T (moving-platform) connects, in parallel, two serial manipulators from Fig. 1 , forming the 5-DOF 3 T 2 R hybrid serial-parallel ( PPP RRR )(P PP RR ) 2-coupled-Cartesian-manipulator of Fig. 2 . Parentheses ( \u00b7 ) enclose the two individual limbs connected in parallel. Five active prismatic P joints, acting along the \u02c6 X An , \u02c6 Y An , \u02c6 Z An , \u02c6 X Cn , \u02c6 Y Cn axes manipulate common-link L T with 5- DOF. Short arrows in schematic In Sc Fig. 2 B identify the five active prismatic P joints. The first active prismatic P joint of the first limb ( P P P RRR ) controls the z W T coordinate of the common-link L T position T W . The same first passive prismatic P joint of the second serial manipulator ( P P P RR ) accommodates changes in distance | z W A 1 \u2212 z W A 2 | , in the \u02c6 Z W W direction, between links L A 1 , L A 2 , as the common-link L T orientation changes. The second prismatic P joint of each limb transmits lateral forces, through links L An , L Cn , to the common-link L T ", + " Possible redundant second active prismatic P 2 joints (not shown), in each limb ( P P 2 P RRR )(P P 2 P RR ) , balance the lateral forces. The active revolute R joint angles \u03b8 A n X , \u03b8 B n Y around the \u02c6 X An , \u02c6 Y Bn axes from Fig. 1 are now passive \u03b8A n X , \u03b8B n Y . Now active prismatic P joints manipulate the orientation of the \u02c6 Z T axis of common- link L T instead of the active revolute R joints Fig. 1 . The two sets of two revolute RR joint axes \u0302 X An \u2226 \u0302 X Cn and \u0302 Y Bn \u2226 \u0302 Y Dn are not geometrically parallel in Fig. 2 . Consequently, as proven below, links L Bn and L Dn rotate, with parasitic-twist-angle \u03b8BD n Z , relative to each other around their common-link \u02c6 Z T axis, as the orientation of common-link L T changes. Therefore, a passive revolute R joint, around the common-link \u02c6 Z T axis, is necessary to accommodate the parasitic-twist-angle \u03b8BD n Z . Figs. 3 , 4 hybrid serial-parallel 5-DOF 2-coupled-Cartesian-manipulator with parallel-revolute-axes. The two sets of two rev- olute RR joint axes \u02c6 X \u2016 \u02c6 X and \u02c6 Y \u2016 \u02c6 Y are geometrically parallel in Figs", + " Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx proven below, since revolute R joint axes \u02c6 Y Bn \u2016 \u02c6 Y Dn are geometrically parallel, there is no parasitic-twist-angle \u03b8BD n Z between links L Bn and L Dn so a passive revolute R joint is not required between them. Therefore links L Bn , L Dn , L T are rigidly connected, forming a single rigid body as shown in drawing In Sm Figs. 3 D, 4 D. The 5-DOF hybrid serial-parallel 2-coupled-Cartesianmanipulator with parallel-revolute-axes of Figs. 3 , 4 have joint topology ( PPP U )(P PP U ) . Note that they have four revolute R joints (two R \u2019s per U ) compared to the non-parallel-revolute-axes configuration Fig. 2 , with a total of five revolute R joints. The manipulator in Fig. 4 is identical to the one in Fig. 3 except that it is rotated \u03b8A 2 Z = \u221290 \u25e6, \u03b8C 2 Z = \u221290 \u25e6 around the \u02c6 Z W axis. The link and coordinate symbols in Fig. 4 are labeled with subscript \u2018 2 \u2019 to distinguish them from the ones in Fig. 3 with subscript \u2018 1 \u2019. Fig. 5 fully parallel 4-DOF 4-coupled-Cartesian-manipulator . Common-link L T in Fig. 5 connects the two 2-coupledCartesian-manipulators from Figs. 3 , 4 . Together they form the 4-DOF fully parallel 4-coupled-Cartesian-manipulator of Fig", + " Positioner-links L An , L Cn . The \u2018positioner-link\u2019 L An , L Cn of each limb { L An , L Bn , L T }, { L Cn , L Dn , L T } rotates around joint axis \u02c6 X An , \u02c6 X Cn located at position A n W , C n W fixed in link L An , L Cn respectively. Two separate Cartesian manipulators independently control linear positions A n W , C n W of positioner-links L An , L Cn in three-dimensions. Additionally, frames F An , F Cn of positionerlinks L An , L Cn may rotate around the \u02c6 Z W axis by angles \u03b8A n Z , \u03b8C n Z respectively. In Fig. 2 for example, frame F An is rotated by \u03b8A n Z = 0 \u25e6 and frame F Cn is rotated by \u03b8C n Z = \u221290 \u25e6. Links { L An , L Bn } form a link-pair with two rotation degrees-of-freedom and links { L Cn , L Dn } form a separate link-pair with two rotation degrees-of-freedom. Coupler-links L Bn , L Dn . The \u2018coupler-link\u2019 L Bn , L Dn of each limb { L An , L Bn , L T }, { L Cn , L Dn , L T } connects to the common-link L T through a revolute R joint or no joint, along the \u02c6 Z T axis, at position T Bn , T Dn relative to frame F Bn , F Dn respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002648_j.mechmachtheory.2020.103891-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002648_j.mechmachtheory.2020.103891-Figure2-1.png", + "caption": "Fig. 2. Spatial 3-PPRR parallel mechanism with mobility F N = 4 and C N = 4 .", + "texts": [ + " Each actuation scheme is then verified through statics analysis. Zhao et al. [12] point out that their method can only provide the number of actuators necessary for the control of the whole mechanism. On the other hand, they cannot guarantee that any valid actuation scheme exists in order to control the end-effector. In other words, the number of actuators needed to control the end-effector can be different from the number needed to control the whole mechanism. Zhao et al. [13] apply this method for a 3-PPRR mechanism, shown in Fig. 2 . The same mechanism is herein analyzed with a novel method, as presented in the next sections and the results compared and validated. The paper is organized follows. Davies\u2019 formulation for degrees of freedom and constraints is first introduced in Section 2 . Based on this approach, a novel condition for the feasibility of an actuation scheme is presented. A novel approach for ac- tuation schemes enumeration, based on matroid theory is then presented in Section 3 . A new algorithm for enumeration of feasible actuation schemes is introduced and applied to some example in Section 4 ", + " For example the basis { d, h }, of dual matroid M \u2217 M N is obtained as: { d, h } = { E \u2212 { a b c e f g }} (20) In the same way all bases of dual matroid M \u2217 M N can be obtained. The sets of Eq. (19) are all feasible actuation schemes for the two-loop planar mechanism of Fig. 1 . Regard that the two- loop planar mechanism is a fractioned mechanism [36] , and each part of the mechanism is a four-bar mechanism, which is controlled by a single actuator. Eq. (19) contains only one actuated joint from each loop, as expected. 4.2. Actuation schemes of a spatial mechanism 3-PPRR The mechanism depicted in Fig. 2 is an overconstrained parallel mechanism. It has been proposed by Gogu [7] and again by Zhao et al. [12,13] . This is, up to the authors\u2019 knowledge, the only mechanism where all actuation schemes are explicitly enumerated. For this reason, this example has the advantage of validating the method proposed herein. The end-effector 4 of the mechanism is connected to the fixed base 0 through three topologically identical PPRR limbs. The two prismatic pairs of each leg, b i and c i with i = 1 , 2 , 3 , are orthogonal to each other", + " The circuit matrix [ B M ] 2,12 is: [ B M ] = [ b 1 c 1 d 1 e 1 b 2 c 2 d 2 e 2 b 3 c 3 d 3 e 3 1 1 1 1 \u22121 \u22121 \u22121 1 0 0 0 0 0 0 0 0 \u22121 \u22121 \u22121 1 1 1 1 \u22121 ] 2 , 12 (21) For the kinematic chain b 1 \u2212 c 1 \u2212 d 1 \u2212 e 1 the motion screws associated with the couplings are: \u02c6 $ m b 1 = [ 0 0 0 1 0 0 ]T \u02c6 $ m c 1 = [ 0 0 0 0 0 1 ]T \u02c6 $ m d 1 = [ 0 1 0 \u2212z d 1 0 x d 1 ]T \u02c6 $ m e 1 = [ 0 1 0 \u2212z e 1 0 x e 1 ]T (22) where x d 1 , z d 1 , x e 1 and z e 1 are the coordinates of couplings respectively d 1 and e 1 in the global coordinate system oxyz shown in Fig. 2 . Notice that the screw systems associated with the other two kinematic chains b 2 \u2212 c 2 \u2212 d 2 \u2212 e 2 and b 3 \u2212 c 3 \u2212 d 3 \u2212 e 3 can be obtained from Eq. (22) by applying a rotation around z -axis of 2 3 \u03c0 and 4 3 \u03c0 respectively. Thus, the following linear transformation can be applied: i $ = [ i T j ] j $ (23) which transforms the coordinates of a screw $ from the coordinates system i into the coordinates system j . For the rotation considered around z axis, the transformation matrix can be written as: [ i T j ] = \u23a1 \u23a2 \u23a2 \u23a3 [ i R j ] 3 , 3 [ 0 ] 3 , 3 [ 0 ] 3 , 3 [ i R j ] 3 , 3 \u23a4 \u23a5 \u23a5 \u23a6 6 , 6 (24) where [ 0 ] 3,3 is the rotation matrix around z -axis of an angle \u03b8 : [ i R j ] = [ cos (\u03b8 ) sin (\u03b8 ) 0 \u2212 sin (\u03b8 ) cos (\u03b8 ) 0 0 0 1 ] 3 , 3 (25) Considering a symmetric configuration of the PPRR mechanism, i", + " In [10] , Qin and Dai investigated the configuration and the actuation of the 2US+UPS asymmetrical parallel mechanism, proposing two actuation schemes. In general, in overconstrained asymmetrical parallel mechanisms not all possible sets of joints can be selected as a valid actuation scheme. Furthermore, not always the actuators near the base are a valid actuation scheme which satisfies Condi- tion 1 . Refaat et al. in [38] introduce four families of three degrees of freedom translational-rotational asymmetrical parallel mechanisms. The mechanism presented in Fig. 3 was originally introduced by Refaat et al. in [38] ( Fig. 2 in the original publication). This mechanism has three degrees of freedom and three redundant constraints. Intuitively, if the revolute joints a and i and the prismatic joint e were actuated and locked, the platform 4 is still capable of rotating along the y axis. Thus, the natural set { a, e, i } is not a valid actuation scheme. On the other hand, if the joints { a, i } and one revolute joints in the set { f, g, h } were actuated, the mechanism is frozen, i.e. Condition 1 is satisfied, and these joints constitute a valid actuation scheme", + " The rank of the dual matroid is r(M \u2217 M N ) = F N , where F N is given by the Modified Grbler-Kutzbach Criterion. The correctness of Algorithm 2 can now be investigated. In Steps 1 \u2212 2 Davies\u2019 method is performed: given a mechanism, matrix [ \u02c6 M N ] can always be obtained [14,35] . In Steps 2 \u2212 3 a linear matroid M and its dual M \u2217 are defined from matrix[ \u02c6 M N ] [27] . Greedy algorithm can be applied to any matroid [27] in Step 5 \u2212 6 . The proposed algorithm can be applied to the 3-PPRR mechanism presented in Fig. 2 . 5.1. Selection of actuation set of 3-PPRR mechanism and of PKM family mechanism First, a set of weights must be defined for the columns of matrix [ \u02c6 M N ] . In general, if the actuators are mounted close to the base of the robot, the total weight of the manipulator is decreased and the power-to-weight ratio is increased [42] . Thus, the following criterion can be employed to establish the coupling weights: \u2022 The actuation set selected must contain the actuators closer to basis 0. If different mechanism specifications are required, different criteria can be stated and different weights sets can be ob- tained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure80.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure80.1-1.png", + "caption": "Fig. 80.1 Designed TPMS-based porous surface models. Isometric view with 3 3 3 unit cells for a primitive surface, b gyroid surface (Zoom in view inside the circle presents the unit cell length (L), pore diameter (D) and wall thickness (T) of each model)", + "texts": [ + " The unit cell was extended periodically in x, y, and z directions by modulating the Eqs. (80.1) and (80.2) as described in Eqs. (80.3) and (80.4) [17]. Primitive \u00f0P\u00de: FP \u00bc cos 2pNx L \u00fe cos 2pNy L \u00fe cos 2pNz L P \u00f080:3\u00de Gyroid:FG x; y; z\u00f0 \u00de \u00bc sin 2pNx L cos 2pNy L \u00fe sin 2pNy L cos 2pNz L \u00fe sin 2pNz L cos 2pNx L P \u00f080:4\u00de where N defines the number of unit cells, L is length of the unit cell, and P is the strut thickness parameter that controls the volume fraction in a unit cell. For a scaffold with 3 3 3 unit cells (Fig. 80.1), value of N was fixed at N = 3, unit cell length L was varied due to which the pore diameter and total scaffold length were altered accordingly as listed in Table 80.1. Seven models for each surface (primitive: P1\u2013P7 and gyroid: G1\u2013G7) type were designed. As a result, a total of fourteen models were designed. P is kept constant at P = 0 to fix the volume fraction at 50%. \u201cSTLWRITE\u201d command was added to the MATLAB code to export the geometry as stl file which can be imported to other software for finite element analysis (FEA)", + " Wall thickness was developed and the model was exported to ANSYS software for FEA after performing surface and volume meshing. The complete procedure is explained in the following subsections. Developing wall thickness and preparing Finite Element Models Stl files generated through MATLAB software were imported in 3-matic software for assigning wall thickness to the surface and carrying out surface and volume meshing. The surface of the models of gyroid and P were thickened with a uniform offset of 102 \u00b5m and 133 \u00b5m, respectively (Fig. 80.1). Since, offsetting the surface disturbed the surface geometry (mesh) of gyroid structures, surface meshing was performed with specific 3-matic tools to improve the quality of surface mesh. Additionally, to reduce the density of the triangles and improve the quality of mesh, quality preserving reduce triangles tool followed with uniform meshing with a target triangle edge length of 0.2 was performed in a manner that element quality is preserved. The shape-quality threshold and maximum geometrical error values were fixed to 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003739_s12239-020-0106-8-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003739_s12239-020-0106-8-Figure5-1.png", + "caption": "Figure 5. Schematic diagram about force analysis of internal structure of modular deformable tire.", + "texts": [ + " The piston compresses the gas very little during the force transmission process, so the force transmission process can be treated as a statics problem. The force transmission equation of the car body load transferred to the tire petal-type connecting block is as follows: (2) where is the pressure provided by the piston fixedly connected with the first petal-shaped connecting block that is subject to the pressure generated by the ground to the gas in the cavity; S is the cross-sectional area of the piston on the petal-shaped connecting block. As shown in Figure 5, the modular deformable tire has six petal-shaped connecting blocks, and the annular cavity and cylindrical cavity inside the tire are connected with each other. When the petal-shaped connecting block in contact with the ground is pressed, the piston fixed to the petal-shaped connecting block will compress the gas in the full gas state in the cavity, while the other five petal-shaped connecting blocks will be subjected to the pressure generated by the compressed gas at the same time. The force transmission equation thus established is as follows: (3) where (i = 1, 2," + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001843_j.ifacol.2019.12.321-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001843_j.ifacol.2019.12.321-Figure1-1.png", + "caption": "Fig. 1. Rigid GFRP hull inspired from Boxfish", + "texts": [ + " Various linear and non-linear control methods have been applied to underwater robots (Yoerger and Slotine, 1985), (Han and Chung, 2014), with slight variations based on the application. In this paper, a nonlinear controller for oscillating fin propelled the underwater vehicle is implemented similar to ' ' the Boxybot (Lachat, Crespi and Ijspeert, 2006), Ostraciiform based robots Micro-autonomous Robotic Ostraciiform (MARCO) (Kodati et al., 2008) and (Wang and Xie, 2014). The vehicle comprised a caudal fin and pair of pectoral fins with a smooth carapace similar to that of a boxfish as shown in Fig. 1. A nonlinear controller is required for the vehicle as to the oscillating angles of the fins results in a nonlinear thrust vector map. Thus, we attempted the back-stepping method for the control design because of this nonlinearity, which increases the complexity in the vehicle control, navigation and guidance due to the nonlinear, time-varying and coupled dynamic behaviour of the subsystems (Kim, 2015). Thus, taking inspiration from some of the works as mentioned earlier, a nonlinear control for an underwater robot whose thrusters can be tilted independently is proposed in this paper", + " A brief discussion on the dynamic modeling of AUV is presented in Section 2 including hydrodynamic and thrust vector terms, followed by the controller design based on backstepping method, to solve the nonlinear problem and derivation of its stability criterion is described in Section 3. Simulations of hovering motion are performed, and the results are displayed in Section 4. Concluding remarks are presented in Section 5. A boxfish carapace inspired Glass Fibre Reinforced Plastic (GFRP) hull is fabricated that can operate until 30 m depth as seen in Fig. 1. The vehicle consists of three rigid fins actuated by three servomotors individually to achieve required thrust force and various motions based on the combination of fin movements. All the fins are attached directly to servo motor shafts to decrease frictional/mechanical losses during their movement. The caudal fin gives the required thrust, and two pectoral fins help change the attitude of the vehicle as well provide additional thrust whenever required. The hull is modular with three compartments with a nose hull freely flooded in the front for mounting oceanographic sensors", + " A moving or vehicle-fixed frame xbybzb is fixed to the center of gravity of the AUV where its movement is given with respect (Fossen, 1994) to the earth-fixed frame xyz. The xbaxis is positive in nose direction of the AUV; yb -axis and zbaxis positive are positive towards the starboard and vehicle the Boxybot (Lachat, Crespi and Ijspeert, 2006), Ostraciiform based robots Micro-autonomous Robotic Ostraciiform (MARCO) (Kodati et al., 2008) and (Wang and Xie, 2014). The vehicle comprised a caudal fin and pair of pectoral fins with a smooth carapace similar to that of a boxfish as shown in Fig. 1. A nonlinear controller is required for the vehicle as to the oscillating angles of the fins results in a nonlinear thrust vector map. Thus, we attempted the back-stepping method for the control design because of this nonlinearity, which increases the complexity in the vehicle control, navigation and guidance due to the nonlinear, time-varying and coupled dynamic behaviour of the subsystems (Kim, 2015). Fig. 1. Rigid GFRP hull inspired from Boxfish Thus, taking inspiration from some of the works as mentioned earlier, a nonlinear control for an underwater robot whose thrusters can be tilted independently is proposed in this paper. A backstepping algorithm is employed for tracking controller to track the desired trajectory in 3D space. The gain optimization was performed using the RMS values of the error. Unlike tilting thruster model proposed in (Jin, Kim, J. Kim, et al., 2015), where a switching controller is used only for two configurations of the tilting angle, we apply control for continuous tilting angle", + " A brief discussion on the dynamic modeling of AUV is presented in Section 2 including hydrodynamic and thrust vector terms, followed by the controller design based on backstepping method, to solve the nonlinear problem and derivation of its stability criterion is described in Section 3. Simulations of hovering motion are performed, and the results are displayed in Section 4. Concluding remarks are presented in Section 5. 2. DYNAMIC MODEL A boxfish carapace inspired Glass Fibre Reinforced Plastic (GFRP) hull is fabricated that can operate until 30 m depth as seen in Fig. 1. The vehicle consists of three rigid fins actuated by three servomotors individually to achieve required thrust force and various motions based on the combination of fin movements. All the fins are attached directly to servo motor shafts to decrease frictional/mechanical losses during their movement. The caudal fin gives the required thrust, and two pectoral fins help change the attitude of the vehicle as well provide additional thrust whenever required. The hull is modular with three compartments with a nose hull freely flooded in the front for mounting oceanographic sensors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001924_icoecs46375.2019.8949914-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001924_icoecs46375.2019.8949914-Figure3-1.png", + "caption": "Fig. 3. Sketch of the motor active part", + "texts": [ + " The presence of working fluid in the air gap leads to an increase in hydraulic losses of more than 100 times. At the same time, the torque of hydraulic losses at a temperature of 20 \u00b0C were 0.012 Nm, which increases the load torque relative to EM without working fluid by 10%; at a working fluid temperature of -30 \u00b0C, the load torque increases by more than 20%. After determining the hydraulic losses and fuel velocities in the EM cavities, electromagnetic calculations of EM with PM were performed. Electromagnetic calculations were performed in Ansys Maxwell software package. Fig. 3 shows the motor active part dimensions obtained from the results of electromagnetic calculations. Fig. 4 shows the winding scheme. The main calculation parameters are shown in Table III. Based on the results of electromagnetic calculations, it can be concluded that the EM design satisfies the technical requirements for the FP with respect to mass, efficiency and rotor speed. Therefore, the next stage of the calculation was the thermal calculations and simulation of processes when one and two phases are closed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003907_012074-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003907_012074-Figure1-1.png", + "caption": "Fig. 1 Lathe assembly model", + "texts": [], + "surrounding_texts": [ + "In this paper, the reducer shaft is analyzed as a machining example. The shaft of the reducer is produced in small batch and is mainly used in the transmission system of the secondary reducer. The key slot 10 is used to fix the transmission gear and the key slot 8 is used to fix the output pulley. The shaft adopts 45 \u00d7 quenching and tempering treatment 230-235hb to meet the high torque requirements of the output shaft. The two-stage reducer is widely used in the production practice, with low price and excellent performance, and has been loved by users. In this paper, the two-stage reducer shaft is taken as the research object to carry out process analysis and three-dimensional modeling, and then its processing is simulated and analyzed by software. Through this example, the importance of virtual simulation of vehicle training is demonstrated. According to different parts, the processing stage is generally divided into rough machining stage, semi finishing stage and finishing stage. In the rough machining stage, simple machining is mainly carried out for the blank, and the rough machining stage is mainly realized by simple rough turning or rough milling. For castings and forgings, the machining allowance is generally about 3mm. Through the semi finishing stage of processing, the surface tolerance level can be achieved, and the surface roughness can be achieved. Finally, through the finishing stage, which is also the final machining of parts, the machining accuracy can be achieved, and the surface roughness can be achieved. Through the reasonable division of the processing stage, we can achieve the best processing of the parts, which can not only improve the production efficiency, but also ensure the quality of the parts processing. This time, it mainly divides the processing stage of reducer shaft, which is mainly used in lathe processing, and the bridge part is mainly processed by washing machine. CISAT 2020 Journal of Physics: Conference Series 1634 (2020) 012074 IOP Publishing doi:10.1088/1742-6596/1634/1/012074" + ] + }, + { + "image_filename": "designv11_80_0003215_012013-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003215_012013-Figure3-1.png", + "caption": "Figure 3. The torsional moment and working boundary conditions of the Cardan shaft", + "texts": [ + "56 WL (1) in which E is the modulus of elasticity, I is the moment of inertia, g is the acceleration due to gravity, L is the length, and W is the weight of the Cardan shaft [1]. The boundary conditions including frictionless support, the fixed support, rotational velocity and moment were used to simulate the same Cardan shaft. The torsional moment is 320 Nm and is applied to all the body with the rotational velocity of 2200 rpm (230.308 rad/sec). The torsional moment and boundary conditions on the Cardan shaft are illustrated in Figure 3. As shown in Figure 4, the mesh used in this study were 381312 nodes and 215054 elements. 10th TSME-International Conference on Mechanical Engineering (TSME-ICoME 2019) IOP Conf. Series: Materials Science and Engineering 886 (2020) 012013 IOP Publishing doi:10.1088/1757-899X/886/1/012013 The results of mode shape and maximum total deformation in four types of materials of Cardan shaft indicate that the highest natural oscillation frequency is in the 10th mode where the Kevlar Cardan shaft has the highest nature oscillation frequency at 1156" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003185_012018-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003185_012018-Figure3-1.png", + "caption": "Figure 3. Four-spring harvester with two by two spaced magnets.", + "texts": [], + "surrounding_texts": [ + "Sources, converting mechanical energy into electrical, such as the electromagnetic harvesters, are increasingly replacing the battery-type low power consuming power-supplying electronic devices [1, 2]. There are usually two types of magnet motion in relation to the coil in these harvesters. In the first case, the magnet moves in parallel to the coil [3] and in the second one - vertically with respect to it, often being located in its air gap [4]. Three different structures of four-spring electromagnetic harvesters are considered here, with two different masses, represented by steel plates with 2 or 4 magnets each and two different coils as well. The harvesters have a fixed coil attached to them, in parallel to which the rare earth magnets move. The aim of the present work is to study the effect of the construction parameters (such as the weight of the concentrated mass, the number of turns, the influence of the coil area, the volume of the magnets, and the size of the air gap) on the output electrical parameters of the studied harvesters. 2. Exposition TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 When applying a sinusoidally varying force with a resonant frequency, the mechanical system: \u201cmass (plate with permanent magnets) \u2013 springs\u201d begins to oscillate in parallel with the coil. Thus the magnetic flux passes differently through the coil and creates an alternating magnetic field in a different way, thus inducing alternating electromotive force. The studied harvesters are nonlinear mechanical oscillating systems. Their simulations made by ANSYS R19.1 take into account the fixture, the gravity effect of the plates with permanent magnets and the mechanical characteristics of the used springs. The horizontal deviation x of the mechanical system \u201cmass (plate with permanent magnets) \u2013 springs\u201d was obtained while modeling the four-spring electromagnetic harvesters, Figure 4. The magnetic field distribution of the three electromagnetic harvesters was obtained by means of FEMM 4.2. Figure 5 shows the magnetic field distribution for the third four-spring harvester with two spaced magnets in two places at zero horizontal deflection, and Figure 7 presents the distribution at maximum deviation. Figure 6 and figure 8 illustrate the magnetic flux density changes along the length of the harvester coil at zero and maximum horizontal deviation. From Figure 6 it can be seen that at zero horizontal deviation the normal magnetic flux density is zero, and at maximum deviation the maximum magnetic flux density Bmax is obtained. TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 The horizontal deviation at a sinusoidally changing force with a resonant angular frequency \u03c9 is also sinusoidal ( ) sinmx t X t\u03c9= (1) The horizontal oscillatory speed is equal to ( ) ( ) sin 2m d x t v t X t d t \u03c0\u03c9 \u03c9 = = + (2) The magnetic flux can be expressed by the change in the cross section of the coil \u0410(t), through which the maximum magnetic flux density Bmax passes ( ) ( ) max dA t \u0424 t B dt = (3) The change in the cross section \u0410(t) over time equals the horizontal deviation variation in the time x(t) along the diameter D\u0441 of the corresponding coil. ( ) ( ) \u0441 dA t D x t dt = (4) From (3) and (4), the magnetic flux through the coil is obtained ( ) ( )max \u0441 \u0424 t B D x t= (5) The induced electromotive force in the coil of the electromagnetic harvester is ( ) ( )d\u0424 t e t N dt = \u2212 (6) From (5) and (6) for the induced electromotive force for no-load mode, it is obtained ( ) ( ) max \u0441 d x t e t N B D dt = \u2212 (7) The amplitude of the induced electromotive force in the coil is a function of the amplitude of the oscillatory speed \u03c9Xm m max \u0441 m\u0415 N B D X\u03c9= (8) The angular frequency of the forced oscillations can be expressed by the frequency f 2 f\u03c9 \u03c0= (9) TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 From (8) and (9) for the amplitude of the induced electromotive force in the coil at a resonance it is obtained active load. In it, Rc and Lc denote the active resistance and the inductance of the electromagnetic harvester coil, \u0435c(t) - the induced electromotive force, and RL is the active load resistance. UL indicates the rectified voltage over the load resistance. The active power, in DC mode, is calculated using the amplitude of the induced no-load electromotive force in the coil at resonance, the voltage on the germanium diode UD and the parameters of the equivalent circuit (14). In the resulting expression, b denotes the attenuation coefficient, which is determined by the logarithmic attenuation decrement \u03b4 [5], where T is the oscillation periodic time and \u0415m(t) is the amplitude of the measured electromotive force 2b m\u03b4= , ( ) ( ) 1 ln m m E t T E t T \u03b4 = + (13)" + ] + }, + { + "image_filename": "designv11_80_0000618_s12541-019-00175-0-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000618_s12541-019-00175-0-Figure5-1.png", + "caption": "Fig. 5 a RB-1 BASE from Robotnik and b NAV350 from SICK, c Robot equipped with a NAV350", + "texts": [ + " Graph-based SLAM (Mref,Xdata,etreshold,etreshold,dthreshold) 1 while(converged); 2 for(all eij); 3 [tij T \u03b8ij] \u2190 ICPOUTLIERREJECTION(Mref,Xdata,etreshold,etreshold,dthreshold); 4 [eij Aij Bij] \u2190 COMPUTEERRORFUNCTION(tij T \u03b8ij); 5 [H] \u2190 COMPUTECONSTRAINT(Aij Bij); 6 [b] \u2190 COMPUTECOEFFICIENTVECTOR(Aij Bij, eij); 7 end for; 8 H11+ = I; 9 [\u2206x] \u2190 SOLVELINEARSYSTEM(H, b); 10 [x] \u2190 UPDATEPOSE(x, \u2206x); 11 end while; 12 return optimal pose; (19) Fij(x\u030c + \u0394x) \u2245 (eij(x\u030c) + Jij\u0394x) T\ud835\udefa(eij(x\u030c) + Jij\u0394x) = eT ij \ud835\udefaeij + 2eT ij \ud835\udefaJij\u0394x + \u0394xTJT ij \ud835\udefaJij\u0394x = cij + 2bij\u0394x + \u0394xTHij\u0394x (20) F(x\u030c + \u0394x) = \ufffd \u27e8i,j\u27e9\u2208C Fij(x\u030c + \u0394x) \u2245 \ufffd \u27e8i,j\u27e9\u2208C cij + 2bij\u0394x + \u0394xTHij\u0394x = c + 2bT\u0394x + \u0394xTH\u0394x (21) Fij \u0394x = 0 (22)H\u0394x = \u2212b (23)x\u2217 = x\u030c + \u0394x\u2217 (24)xT i = ( tT i , i ) The error function in line 4 is given by (26), and Aij and Bij are the Jacobian matrices of the error function, which are given by (27). The matrix H and vector b are correlated in (22) as well as in lines 5 and 6. The process of optimizing H and b is shown in (23) as well as in lines 8 and 9. Figure\u00a05a shows the RB-1 BASE from Robotnik and Fig.\u00a05b shows the scan sensor NAV350 from SICK. The robot used in this experiment is RB-1 BASE from Robotnik [22], which has a size of 500\u00a0mm \u00d7 500\u00a0mm \u00d7 251\u00a0mm. It features a load capacity of 50\u00a0kg, a 4th generation Intel i7 processor and 8\u00a0GB RAM. It has a maximum speed of 1.5\u00a0m/s and can measure odometry data using an encoder sensor equipped on the robot. The scan sensor is a laser rangefinder (LRF), NAV350 [23] from SICK, which has a maximum range of 250\u00a0m and 360\u00b0 with a resolution of 0.25\u00b0. This experiment was performed using the robot, RB-1 BASE, equipped with a scan sensor, NAV350 (Fig.\u00a05c), and the range data for the given environment was obtained using this robot. The experimental environment is shown in Fig.\u00a06. The raw odometry and range data were obtained from the encoder sensor and LRF, respectively. The odometry data, which are obtained from the encoder as raw data, are used for the initial guess, and they are transformed as the scan matching process is carried out with the ICP or ICP Outlier Rejection algorithm. The loop closing process is defined as the recognition of the robot (25)zT i = ( tT ij , ij ) (26)eij(x) = ( RT ij (RT i ( tj \u2212 ti ) \u2212 tij j \u2212 i \u2212 ij ) (27) Aij = eij(x) xi = ( \u2212RT ij RT i RT ij RT i i ( tj \u2212 ti ) 0 \u22121 ) Bij = eij(x) xj = ( RT ij RT i 0 0 1 ) 1 3 returning to the same position from where it started" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002614_j.promfg.2020.04.235-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002614_j.promfg.2020.04.235-Figure1-1.png", + "caption": "Fig. 1. Process-chain of the Tailored Forming Technology for a hybrid bevel gear", + "texts": [ + " The Tailored Forming Technology offers a promising method for manufacturing of high-performance components with locally specific characteristics adapted to the particular application area. In [8,9], Behrens et al. present challenges arising due to different material properties and offer possible solutions for heating and forming strategies. Within this article, the complex process chain for the production of a hybrid bevel gear is presented. The following process steps for the production of hybrid bevel gear based on the Tailored Forming Technology are required (cf. Fig. 1). At first, the different layers are applied on the cylindrical semifinished product by means of deposition welding. This is one of the most frequently used process in joining technology since it is preferred for the use of expensive functional materials on low-priced base materials [10]. Within a subsequent hot forging process, the semi-finished product is formed to a hybrid component with a complex gear geometry. Additional processing steps such as heat treatment and machining finalise the process chain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001511_tmag.2019.2942509-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001511_tmag.2019.2942509-Figure3-1.png", + "caption": "Fig. 3. Finite-element mesh in the silicon core iron.", + "texts": [ + " A time-harmonic analysis was performed at 60 Hz. The stator and rotor silicon steel cores have a conductivity of \u03c3 = 2.22 \u00d7 106 S/m and a relative permeability of \u03bc\u22a5 = 4000. The coil exits the stator at a distance R = 0.053 m. Other dimensions are w = 0.004 m, \u03b8 = 0.075 rad, and h = 0.050 m. A current of 120 A flows in the coil. The model was enclosed with an air cylinder with a relative permeability of \u03bcr = 1, a radius of 0.2 m, and a height of 0.2 m. Finite elements of 130 000 were employed in the model. Fig. 3 shows the mesh used to simulate the magnetic field distribution in the core. Eighteen layers of finite elements of a thickness 0.689 mm were employed to properly account for the skin-effect. Nine layers were employed in the upper and nine in lower part of the machine. The steps that have been employed for the numerical computation of the leakage inductance are as follows. 1) First, the electromagnetic field in and outside the machine was calculated using FEM. 2) Second, the electromagnetic field energy accumulated outside the machine, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003882_0142331220962793-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003882_0142331220962793-Figure1-1.png", + "caption": "Figure 1. Geometry of three-dimensional tracking problem.", + "texts": [ + " Moreover, in the presence of interference, two-sensor information fusion combined with guidance laws are proposed to enhance the precision of trajectory tracking. The remainder of this paper is organized as follows. The kinematics equations of the pursuer and the target are derived. Then, the guidance laws of the deviated pursuit and the proportional navigation are discussed. This paper proposes the control strategy of tracking under guidance laws. This paper presents two-sensor information fusion enhancement. A series of simulation results are given and then some conclusions are drawn. Kinematics models of pursuer and target As shown in Figure 1, (xP, yP, zP) denote the coordinates of the pursuer in the three-dimensional coordinate system, vP is the linear velocity, uP and fP are the flight path and heading angles. (xT , yT , zT ) denote the coordinates of the target, vT is the linear velocity, uT and fT are the flight path and heading angles. The geometry of three-dimensional tracking problem is illustrated in Figure 1. PT denotes the line of sight between the pursuer and the target. s is the pitch angle of PT , and g is the yaw angle of PT . The relative distance between the pursuer and the target is given by r = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi (xT xP) 2 +(yT yP) 2 +(zT zP) 2 q : \u00f01\u00de The pitch angle of PT and the yaw angle of PT are given by tans= zT zPffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi (xT xP) 2 +(yT yP) 2 q , \u00f02\u00de tang = yT yP zT zP : \u00f03\u00de In this paper, the polar representation is applied to derive the kinematics equations between the pursuer and the target in the three-dimensional space" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001429_j.mechmachtheory.2019.103674-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001429_j.mechmachtheory.2019.103674-Figure6-1.png", + "caption": "Fig. 6. Exemplary four-bar mechanism as virtual prototype with limit positions and supporting spring.", + "texts": [ + " Sensors measure the current position or force values as input for the multi-body simulation of the linkage. For the purpose of haptic synthesis a model of the virtual prototype, i.e. the to be designed mechanism, has to be simulated online and in real-time using the PLC. This section discusses different aspects of modelling and simulating the multibody dynamics of mechanisms under these conditions. As an example, a simple four-bar mechanism with limit positions and a spring as force support is used, see Fig. 6 . The moving plane is perpendicular to gravity, whereby related terms can be ignored. For the derivation of the dynamics equation an analogous approach to the RePlaLink model, compare Section 3 , is followed. The handle pose x H, VP can be calculated with the direct kinematics based on the crank angle \u03c3 and the geometric parameters of the four-bar p VP , see (14) , and again vice versa the inverse kinematics with (15) . x H, VP = DK VP ( \u03c3, p VP ) = [ x H, VP y H, VP \u03d5 H, VP ]T (14) \u03c3 = IK VP ( x H, VP , p VP ) (15) For this one degree of freedom case the Jacobian can be viewed as a transmission ratio I and similarly for the centres of gravity for the translations I T and the rotations I R ", + " In this configuration, a change of the branch, in which the mechanism is assembled, is possible. This singularity is only caused by the mathematical formulation of the direct kinematics and is not a property of the hand-actuated mechanism itself. However, it has to be prevented. The assemblability constraint is therefore not only applied to mechanisms that have a crank with limited range of motion but to every mechanism. For a hand-actuated mechanism such as considered here, the actual actuation takes place at the handle mounted on the coupler (see Fig. 6 ). Regarding this, two singularities may occur. The first one occurs when all links of the four-bar mechanism are in-line. In this case, the motion of the mechanism is not unique with respect to the motion of the handle. This is already prevented by the extended assemblability constraint. The second type of singularity occurs if the position of the handle equals the instantaneous centre of velocity of the coupler. Then, the mechanism cannot not be moved by applying forces at the handle since any forces are fully compensated by constraint forces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.9-1.png", + "caption": "Figure 1.9. Rotation of the vector u around the vector n; Left: perspective view; right: view from above", + "texts": [ + " 6) Using the previous questions, give the expression of a rotation around the vector ! = (1, 0, 0) of angle \". 7) Write an expression of the Euler matrix that uses Rodrigues\u2019 formula. EXERCISE 1.5.\u2013 Geometric approach to Rodrigues\u2019 formula Let us consider the rotation Rn,% of angle & around the unit vector n. Let u be a vector that we will subject to this rotation. The vector u can be decomposed as follows: u =< u,n > \u00b7n2 34 5 u|| + u\" < u,n > \u00b7n2 34 5 u\" where u|| is collinear to n and u% is in the plane P% orthogonal to n (see Figure 1.9). 1) Prove Rodrigues\u2019 formula given by: Rn,% (u) =< u,n > \u00b7n+ (cos&) (u\" < u,n > \u00b7n) + (sin&) (n ' u) . 2) Using the double cross product formula a' (b ' c) = / aTc 0 \u00b7 b \" / aTb 0 \u00b7 c, on the element n' (n ' u), show that Rodrigues\u2019 formula can also been written as: Rn,% (u) = u+ (1\" cos&) (n' (n ' u)) + (sin&) (n ' u) . Deduce from this that the matrix associated with the linear operator Rn,% is written as: Rn,% = $ % 1 0 0 0 1 0 0 0 1 & ' +(1\" cos&) $ % \"n2 y \" n2 z nxny nxnz nxny \"n2 x \" n2 z nynz nxnz nynz \"n2 x \" n2 y & ' +(sin&) $ % 0 \"nz ny nz 0 \"nx \"ny nx 0 & ' " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002500_icitisee48480.2019.9003796-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002500_icitisee48480.2019.9003796-Figure3-1.png", + "caption": "Fig. 3. Initial models with different stator tip", + "texts": [ + " Parallel Tip is an initial model with a stator tooth tip complement to the inner stator radii \u2022 Processing Stage This stage is the performing of the simulation process to obtain the cogging torque value and reference generator density flux and model modification. Simulations were carried out with the support of FEMM software version 4.2. The models are simulated with the mechanical displacement angle of the rotor every 1 degree. For periods of cogging torque waves ranging from 0 degrees to 15 degrees. Before being simulated, all four models are split into mesh with the various number of mesh based on the design of each model. The mesh created in each model can be viewed in the fig. 3 below: 45443 Nodes 56249 Nodes (a) No Tip (b) Tiered Tip Fig. 3 (a) is a no tip generator model with a mesh set up by since many as 45,443 nodes. The mesh in Fig. 3 (b) is a tiered tip made by 56,249 nodes. The mesh in the Round tip model in Fig. 3 (c) is set up by 56,495 nodes, while the parallel tip model in Fig. 3 (d) is made by 50,638 nodes. The quantity of meshes composed by nodes is influenced by the complexity of the design conceived in the generator model. \u2022 Post-Processing Stage / Solution After the simulation disposing of, the cogging torque value is attained. Cogging torque waves are gathered for each modified model and reference model. From these proceeds, we can identify the greatest wave peak and the least wave peak gathered from each model. Formerly, the density flux distribution from each model is analyzed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003807_icuas48674.2020.9214047-Figure18-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003807_icuas48674.2020.9214047-Figure18-1.png", + "caption": "Figure 18. 3D CAD Design of the aircraft", + "texts": [ + " To ensure this, a tilting mechanism has been designed, necessary strength analysis has been made and the mechanism has been manufactured. The mechanism is given in Figure 17. Necessary tests are conducted for the thrust vectoring mechanism. It is seen that the vectoring mechanism successfully tilt the engine to the desired angle in expected time without overshoot. 880 Authorized licensed use limited to: Newcastle University. Downloaded on October 18,2020 at 15:29:54 UTC from IEEE Xplore. Restrictions apply. C. 3D Design of the Proposed Aircraft The aircraft is designed and 3D drawings are prepared. Figure 18 shows the 3D CAD model of the jet powered VTOL aircraft. There are four jet engine tilting gimbal mechanism made from aluminum. Fuel needed to power the engines is stored in four 12 liter fuel tanks, colored in orange in the same figure. The manufacturing process of the aircraft is still going on. An image of the aircraft is given in Figure 19. The tilting mechanisms and the aircraft frame are manufactured from aluminum. We would like to thank The Scientific and Technological Research Council of Turkey (TUBITAK) for funding project 118E192" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002901_012017-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002901_012017-Figure1-1.png", + "caption": "Figure 1. Vehicle axis system ISO 8855-2011", + "texts": [], + "surrounding_texts": [ + "It is important to study: maneuverability by following the speed of movement and lateral stability by minimizing the lateral slip angle. The four-wheel vehicle model was subsequently examined in the case of a low-level direction for ESP. The clutches were not controlled to a high level of control since two different vehicle models were considered. In addition, no lateral speed adjustment was provided, whereas a vehicle equipped with active front and rear steering systems (AFS and ARS) can allow lateral transient movement, for example to avoid an obstacle. The current state shows that we are using complex and reliable controllers based on unrelated simplified vehicle models, or that we are separating a complex vehicle model to use simplified controllers in each direction. The authors do not give priority to either the first or the second approach. The new approach has been sufficiently studied, in which a relatively sophisticated and robust high level controller is used, based on a relatively complex model of four-wheeled vehicle with an optimal coordination strategy. The objective is to evaluate dynamic coupling at the vehicle level to justify the structure of the high-level controller. A twin longitudinal transverse control requires a paired longitudinal transverse vehicle model. Future autonomous vehicles (AV) must simultaneously drive longitudinal, lateral and, ultimately, vertical controls. These speakers are connected because they are connected to the same system. For global chassis control (GCC), the vehicle\u2019s internal dynamic clutches must be considered. These connections may be more restrictive or, conversely, more relaxed depending on the chassis systems integrated in the same vehicle. For example, with two-wheel steering (2WS), a vehicle can avoid an obstacle only by changing its course (angle of movement). Unlike the vehicle 2WS, vehicle 4WS, four-wheel steering in one direction, you can bypass an obstacle without changing its direction. Since it is impossible to know the future hardware design of cars, and since the design of cars may differ from different manufacturers, the authors of the article are developing a new detailed global car simulation. In order to replace the driver and make cars autonomous, it is necessary to introduce additional embedded systems. In addition, in order to distinguish a car brand from competitors, various subsystems can be implemented by various manufacturers. One thing is certain, future vehicles can be heavily overloaded. A layered architecture is a good option for future automotive control tasks. This architecture provides, in particular, the following criteria: adaptability, fault tolerance, dynamic reconfiguration, extensibility and modularity. The car is equipped with an Active Rear Steering (ARS) system, a vehicle-based braking dynamics control system (VDC) and two rear wheel electric motors for vectorizing rear torque (RTV). The generalized efforts required to move the vehicle can be optimally distributed using optimization strategies based on optimization (CA) on four buses. Then, such bus forces can be converted into actuator commands and activate the corresponding system, avoiding any internal conflicts. DS ART 2019 IOP Conf. Series: Materials Science and Engineering 832 (2020) 012017 IOP Publishing doi:10.1088/1757-899X/832/1/012017 DS ART 2019 IOP Conf. Series: Materials Science and Engineering 832 (2020) 012017 IOP Publishing doi:10.1088/1757-899X/832/1/012017 There are three types of control: \u2022 High level control The goal here is to calculate the required forces on the vehicle propeller to track the desired speeds. Dynamics at this point is characterized by inertial parameters such as mass and moment of inertia. These parameters are usually uncertain, which requires a certain degree of reliability. A multi-input multi-output (MIMO) controller is required to account for various connections. Since the vehicle is equipped with ARS, VDC and RTV and access to active suspensions is not permitted, only a flat vehicle model can be considered at a high level. However, when evaluating tire potential, the importance of vertical dynamics should be considered. \u2211=Ss+Suf+Sur \u03a3: total body mass M and center of gravity (CoG) G, Ss: spring-loaded mass Ms and CoG Gs, Suf: front unsprung mass Muf and CoG Guf, Sur: rear unsprung mass Mur and CoG Gur. \u2022 Medium level control [8]. This intermediate level aims at coordinating the wheel systems in order to avoid differences and to generate high level general forces. To do this, all the forces must be distributed optimally on the four tires in order to activate the system with the desired level of effort. Since the number of forces on the tires exceeds the total forces that must be applied to the propeller of the vehicle, the system is overactivated, which leads us to the CA problem." + ] + }, + { + "image_filename": "designv11_80_0002301_icisct47635.2019.9011865-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002301_icisct47635.2019.9011865-Figure1-1.png", + "caption": "Fig 1. Dynamic Simulation.", + "texts": [], + "surrounding_texts": [ + "Index Terms\u2014 Dynamic simulation, autodesk inventor professional, demo platform, kinematic pairs.\nI. INTRODUCTION\nAt present, the simulation of dynamic characteristics of mechanisms and actuators of machines for various purposes is quite an urgent task for the designers. It allows them to determine how effectively the designed machine will function in real conditions and take, if necessary, corrective action to ensure correct operation, strength and durability of its mechanisms and components [1].\nII. FORMULATION OF THE PROBLEM\nIn design when realizing the project, the Autodesk Inventor Professional 11 software package family is preferred. It is the only software product today that allows us to switch the modeling process from 2D to 3D and vice versa. The Dynamic Simulation module, entering, among other innovations, this software allows us to efficiently calculate any dynamic systems, that is, to design and analyze the operation of the main types of existing mechanisms, and any combination of them.\nIII. DECISION\nThe Dynamic Simulation module implements the latest technologies for modeling the dynamic operation of mechanisms. It allows us to analyze the mechanism structure; to apply external forces and moments to the nodes of the mechanism; to build the trajectories of the nodes points motion; to display the graphs of velocities and accelerations, calculated by simulating the movement of the mechanism\nparts at various stages of operating cycle; to calculate the forces and moments that occur in the mechanism nodes in operation process; to record animated clips of mechanism motion. The module also allows us to determine the working cycle time, to reveal \u201charmful\u201d intersections, to calculate dynamic loads in detail [2].\nAt the final stage, the simulation of the mechanisms operation will help to quickly solve such typical designer\u2019s tasks as: choosing the drives and moving units position, determining the characteristics of springs and dampers, determining the mass parameters, the radii of the rollers, cam profiles, etc. Using a virtual prototype of the product and the opportunities of modeling dynamic operation of mechanisms, the designer gets the opportunity to consider and analyze a large number of product configurations with various parameters in order to choose the most optimal option. When simulating the dynamics of mechanisms in Autodesk Inventor, one needs to know that:\nThe simulation process of dynamic operation of mechanisms in Dynamic Simulation occurs according to the following scheme: 1. In the standard Autodesk Inventor working environment,\n\u201chard\u201d subassemblies are created, that is, the groups of\nparts that will move relative to each other [3]. 2. Dynamic Simulation sets the types of connection of the\nsubassemblies to each other from the existing list (rotation, sliding, rolling, various contact interactions, springs, pushers, etc.).\n3. Connection parameters are determined (gravity, friction, damping, superimposed movements and various external forces).\n4. The process of operation simulation is started - the time of operation and the time for each step are set.\n5. The analysis of the results is carried out - the positions of the parts, velocities, accelerations, reactive forces and torques, driving forces and their moments are determined.", + "Dynamic Simulation module gives an extensive set of kinematic pairs (Fig 2). It provides both standard types of kinematic pairs and various special types that describe the operation of gears and worm-gears with moving and fixed axles, belt and chain gears, cam mechanisms, ratchet and pinion gears, and set elastic ties and three-dimensional contact between the bodies.\nKinematic pairs are set by two methods.\nThe first method is the creation of kinematic pairs by setting the degrees of freedom (Fig 3). To do this, select the desired type of kinematic pair from the menu, select the parts that form the kinematic pair and then connect them together, defining edges, axes, planes, points, etc. of interacting components. After the end of operation, the parts are automatically set to the specified location [5].\nThe second method is to use assembly dependencies created in the context of the Autodesk Inventor assembly (Figure 4). To do this, select two parts that form a kinematic pair, and activate the existing assembly dependencies\nFor the model to work, it is necessary to set to its nodes the appropriate movements. For this, we can operate with various degrees of freedom of kinematic pairs, setting the necessary displacements, velocities, and accelerations [6].\nWhen all kinematic pairs of the mechanism are created and the movements set, we can view and edit the structure of the mechanism by clicking the Repair Redundancies window (Fig 5).", + "IV. RESULT\nHere are all kinematic chains of the mechanism, the number of degrees of freedom, and the number of static indefinability \u043c\u0435\u0445\u0430\u043d\u0438\u0437\u043c\u0430 of the mechanism. Using the Highlight chain\u2019s components button, we can highlight on the screen the links that enter some sort of kinematic chain; the Test button allows us to check the kinematics of the mechanism operation.\nNow proceed to analyzing the correct functioning of the mechanism, that is, to obtaining graphic data on its operation, for which it is necessary to analyze the forces, moments, velocities, accelerations and displacements acting in the model.\nTo do this, use the Output Grapher window. In our case, we want to determine the torque on the motor shaft along the X and Y axes. By double-clicking on the plot point of interest, we can see, for example, the maximum values of the Y-axis torque in N\u0445mm (Fig. 6). In this example, we note that the Dynamic Simulation module in Autodesk Inventor Professional 11 will provide indispensable assistance in the design of such mechanisms as rotor lines, piston devices, mechanical manipulators, and many others, in design of which the use of dynamic modeling methods is necessary.\nV. CONCLUSION\nThe developed and elaborated technological demo platform complies with the requirements and operational data. Developed models will be a convenient tool for the designers and a powerful analytical resource. The solutions found will allow not only to store and integrate data, but also to display the process of objects operation in 3D models [7].\nThe 3D information model serves as a three-dimensional interface for accessing data - the user gets the opportunity to view the information he needs by selecting and clicking on the corresponding model element. A patent application has been filed for this product.\n[1] Yakubov M.S., Kiriakidi A.S., \u201cModern computer technology and the development of visualization systems in Uzbekistan,\u201d Science of the 21st century: questions, hypotheses, answers, vol. 2 pp.110-112, 2014.\n[2] Yakubov M.S., Kiriadi A.S., Designing 3D models of scientific and technical developments, Reports of a scientific and practical conference, 2014. pp. 173-174.\n[3] Yakubov M., Varisov \u0410., \u201cSome issues of citizens' access to information resources in e-government system,\u201d Academy of Public Administration under the President of the Republic of Uzbekistan Society and Governance, vol. 2, 2014, pp. 64-74.\n[4] J. Lee, B. Ware, \u201cThree-dimensional graphics and animation,\u201d 2nd ed., Williams, 2002, 640 p.\n[5] D. Hearn, M.P. Baker, \u201cComputer graphics and OpenGL standard,\u201d 3rd ed., 2005, 1168 p.\n[6] E. Angel, \u201cInteractive computer graphics. Introductory course based on OpenGL,\u201d 2nd ed., Williams, 2001, 592 p.\n[7] G. Snook, \u201cReal-time 3D landscapes in C ++ and DirectX 9,\u201d 2nd ed., Kudits-press, 2007, 368 p. - ISBN 5-9579-0090-7." + ] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure6.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure6.2-1.png", + "caption": "Fig. 6.2 Stage 1 of the proposed joining method", + "texts": [ + " By using the simple die which is used in case of single tube, the locking of tubes did not take place. So a die having groove is designed to fabricate the joint. The entire process can be divided into three stages, i.e., stage 1, stage 2 and stage 3. Stage 1: As both the tubes are of different length, i.e., internal tube is larger than the external tube, when axial compression takes place the internal tube starts to deform at the die surface. Since there is no contact between the punch and external tube, it rests at the bottom portion without having any deformation which is shown in Fig. 6.2. 68 E. Premananda and R. Ganesh Narayanan Stage 2: In this stage, the external tube comes in contact with the punch and it also starts deforming above the internal tube. By the time, the internal tube keeps on deforming and acquires the shape of groove of the die. The process of stage 2 is shown in Fig. 6.3. Stage 3: As the upper portion of the die groove is heading inward, the leading edge of internal tube starts deforming in the inward direction. After a particular stage, the tubes get mechanically locked as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002992_icusai47366.2019.9124797-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002992_icusai47366.2019.9124797-Figure2-1.png", + "caption": "Fig. 2. Systems of coordinates", + "texts": [ + " 1 gives a helicopter attitude simulation system, which is a simplification of the common helicopter system. It removes the position motion of the helicopter and only contains the attitude movement. The multi-input- multioutput, non-linear and strong coupling characteristics in the attitude motion are fully preserved. The platform can be used for simulation and experiment of helicopter attitude control conveniently. First of all, two reference frames are defined to describe the attitude of the helicopter system. As shown in Fig. 2, one is the fixed frame binding the pedestal, which is defined by OXYZ, where O is the intersection of pitch axis and yaw axis of the experimental platform, the X-axis is perpendicular to yaw axis towards initial direction, the Z-axis and the yaw axis coincide and point downward, and the Y-axis points to the right to form a Cartesian coordinate system. The other one is the body frame ObXbYbZb, where Ob is the intersection of the short rod and the long rod, the Xb-axis coincides with the long rod and points to the motor side, the Yb-axis coincides with the short rod and points to the right motor, and the Zb-axis perpendicular to the plane of XbObYb and points downwards" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002158_cac48633.2019.8997353-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002158_cac48633.2019.8997353-Figure6-1.png", + "caption": "Figure 6", + "texts": [ + " Tilting the body causes the robot's COM projection to approach one of the stable boundaries. So the purpose of our control is to control the length of the three virtual straight legs so that the body of the robot is parallel to the horizontal plane. The main idea of the reference trajectory design is that when the platform is tilted, there is a height difference at the end of the three legs. Although the heights of the three virtual leg support points are different, by extending the virtual leg, the three connection points on the body can be at the same height [14]. As shown in Figure 6, F1, F4, and F7 are the coordinates of the three connection points on the fuselage. F1s, F4s and F7s are the coordinates of the support points of the three virtual legs. The geometric center of the triangle consisting of three support points is used as the origin to establish a Cartesian coordinate system. The distance from the geometric center to each point is L. The dotted line refers to the robot's virtual straight leg, and the actual robot leg has joints. Form as (7),lb1 is the length of the virtual leg 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001812_phm-qingdao46334.2019.8943043-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001812_phm-qingdao46334.2019.8943043-Figure2-1.png", + "caption": "Figure 2. Bearing testbed (a) and failure bearing with outer race wear (b)[7]", + "texts": [ + " ANALYSIS PROCESS AND CASE This paper selects the bearing data provided by the Institute of Design Science and Basic Components at Xi'an Jiaotong University [7] as input. This data is obtained from several accelerated degradation tests of the normal operation to the severe failure of rolling bearing. During the acceleration, it is considered as severe failure once the amplitude of vibration signal reaches 10 times the maximum amplitude ( maxA ) at the normal operation phase of the bearing. The bearing model is LDK UER204, and its bearing test bench is shown in Figure 2. The test conditions are changed by changing the rotational speed and radial force. The bearing with machining condition of 35 Hz/12 kN and the fault type of outer ring wear (Figure 2) is selected as the analysis object. The specific parameters are shown in Table I, and the original signal is shown in Figure 3. 1) Time domain analysis. Time domain analysis is the analysis of data directly in the time domain. The time domain signal indicatores are mainly used to judge the hidden fault, failure degree and its development trend of equipment, which is featured by intuitiveness and accuracy. In this paper, 14 commonly used time domain indicatores are adopted for data analysis, including mean value, absolute mean value, root mean square value (RMS), amplitude of RMS, peak-to-peak value, peak factor, standard deviation, kurtosis, kurtosis factor, form factor, pulse factor, margin factor, skewness and skewness factor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000044_scems.2018.8624865-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000044_scems.2018.8624865-Figure1-1.png", + "caption": "Fig. 1 The selection area of voltage vector", + "texts": [ + " The predictive control selects the optimal voltage vector from aggregate of voltage vectors with different angle. More selectable voltage vectors will lead to better control performance, but it will cause heavier calculation burden. In this paper, control performances of surface PMSM DTC system using different aggregates of voltage vectors are studied. Considering control performance and calculation burden, an ideal aggregate of voltage vector is given. VOLTAGE VECTORS WITH VARIABLE ANGLE In stator flux reference frame, selection area of the DTC in surface PMSM DTC is shown in Fig. 1 [8]. The selection area of (increase stator flux and torque) is (0\u00b0, 90\u00b0); the selection area of (decrease stator flux and increase torque) is (90\u00b0, 180\u00b0 \u2212 ); the selection area of (decrease stator flux and torque) is (180\u00b0, 270\u00b0) ; the selection area of (increase stator flux and decrease torque) is(270\u00b0, 360\u00b0 \u2212 ), where is torque angle. Neglecting voltage drop due to stator resistance, after applying voltage vector for t\u0394 , the change of stator flux is shown in Fig. 2, where \u03b1 is the angle between the applying voltage vector and stator flux" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003983_icma49215.2020.9233584-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003983_icma49215.2020.9233584-Figure2-1.png", + "caption": "Fig. 2 The model for a single-stage planetary gearbox", + "texts": [ + " For instance, planet gear and carrier are connected by alignment pin, bearings, gaskets and other parts, here it is simplified to a cylinder. Secondly, using ABAQUS software to create the flexible body models of inner ring gear and planet carrier, as shown in Fig. 1. Finally, the rigid body model is imported into ADAMS software, then constructed flexible body models of inner ring gear and planet carrier are imported to replace corresponding parts of the rigid model. The rigid-flexible coupling model is shown in Fig. 2. When gears are meshed with each other to transmit torque, strain will occur at teeth root of gear. Hence, the tooth root strain signals contain important information of gear meshing process, which can be used as the basis to judge whether the model parameter setting is reasonable or not. The strain signal form is the key to verify constraints of flexible body models and mesh settings. If the mesh and constraints settings are unreasonable, the constructed rigid-flexible coupling model will be quite 1964 Authorized licensed use limited to: Carleton University" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001727_icems.2019.8921440-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001727_icems.2019.8921440-Figure2-1.png", + "caption": "Fig. 2. Voltage vector and voltage limit", + "texts": [ + " The inverter maximal phase voltage limit Ulim is defined by the dc-bus voltage Vdc for the linear modulation mode, whereas the current limit Ilim is defined by the available output current of the inverter where the following must be fulfilled: 2 2 lims d qu u u U= + \u2264 (6) 2 2 lims d qi i i I= + \u2264 (7) where, us is the voltage vector magnitude; is is the stator current vector magnitude. According (5) and (6), (8) are deduced when the IPMSM operates in steady state and the resistance voltage drop can be neglected in high speed. 2 2 2( ) ( )\u03c8 \u03c9 + + \u2264 lim d d f q q e U L i L i (8) When space vector pulse width modulation (SVPWM) strategy is applied, the voltage limit is / 3dcV , as shown in Fig. 2. Authorized licensed use limited to: University of Canberra. Downloaded on June 07,2020 at 02:03:39 UTC from IEEE Xplore. Restrictions apply. The block diagram of IPMSM drive system is shown in Fig. 3. The voltage source inverter is used to drive the PMSM. The switching states of the six switches are controlled by the SVPWM strategy. IV. FLUX-WEAKENING CONTROL FOR IPMSM The full operation range of IPMSM can be divided into the low speed area, which is below the motor base speed, and the constant power area, which is above the base speed, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000029_s12239-019-0013-z-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000029_s12239-019-0013-z-Figure5-1.png", + "caption": "Figure 5. McPherson strut type (RSSS-SC) suspension with wrenches acting on the wheel hub.", + "texts": [ + " The wheel hub is connected to the vehicle body by two R-S links, each of which exerts two zero-pitch force wrenches on the wheel hub acting at the center of the S joint as shown in Figure 1 (c), and by an S-S link which has a zero-pitch force wrench acts through the centers of two S joints as shown in Figure 1 (e). These five wrenches are independent, and the twist reciprocal to the five wrenches determines the instant screw axis of the wheel with respect to the vehicle body for the bump-rebound motion and its pitch. Figure 5 shows an RSSS-SC spatial mechanism which is equivalent to the McPherson strut type suspension where the wheel hub is connected by an S-C link which exerts two force wrenches acting perpendicular to the S-C link at the S joint, by an S-S link with a zero-pitch force wrench acting along the S-S link, and by an R-S link with two zero-pitch force wrenches. Hence, the five wrenches acting on the wheel hub is determined. Figure 6 shows a 5-SS multi-link type suspension mechanism. Each S-S link exerts a zero-pitch force wrench along the line connecting the centers of the S joints, and they form five wrenches acting on the wheel hub" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000755_b978-0-08-102211-5.00004-8-Figure4.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000755_b978-0-08-102211-5.00004-8-Figure4.2-1.png", + "caption": "Figure 4.2 Manufacturing process of dry spinning.", + "texts": [ + " There are two types of solution spinning: (1) dry spinning in which polymers are dissolved in a highly volatile solvent and the fibres are formed by evaporating the solvent during spinning and (2) wet spinning in which the solvent is unable to be evaporated from the solution and the fibres have to be spun in a nonsolvent bath to remove the solvent and to coagulate the fibres. The dry spinning is used to produce acrylic fibre, acetate fibre and polyvinyl chloride fibre from acrylonitrile, cellulose acetate and polyvinyl chloride, respectively. Fig. 4.2 schematically shows the basic manufacturing process of dry spinning. The polymer is dissolved in the proper volatile solvent, such as ether and acetone, to form the polymer solution first. Then this polymer solution is mixed with additives and filtered to form a low viscosity dope. Next, the dope is further filtered, de-aired and preheated to maintain the quality of the fibres produced before being added into the extruder. After the dope is prepared, it is added into the extruder and pumped through the filters of the extruder to achieve the required consistency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001288_j.matpr.2019.08.229-Figure14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001288_j.matpr.2019.08.229-Figure14-1.png", + "caption": "Fig. 14. Results of longitudinal stress contours(a) Aluminum (Al), (b) Copper (Cu), and (c) Structural steel materials.", + "texts": [], + "surrounding_texts": [ + "Fig. 12. Results of equivalent stress contours of alloys (a) Aluminum (Al), (b) Copper (Cu), and (c) Structural steel materials.\nFig. 13. Results of hoop stress contours of alloys (a) Aluminum (Al), (b) Copper (Cu), and (c) Structural steel materials.\nTable 3 Comparison of unidirectional epoxy carbon and aluminum alloy cylinders.\nS.No. Material Density (kg/m3) Thickness (mm) Diameter (mm) Volume (m3) Mass (kg)\n1. UD epoxy carbon (230 GPa) Prepreg 1490 101.54 1879 7.0248 10,467 2. Aluminum Alloy 2712 101.54 1879 7.1751 19,459\n6. Conclusions\nIt is concluded that, composite materials will have high strength by comparing with all materials and alloys. For 14 MPa, internal pressures of CNG Auto applications, Composite materials are the best choice. It is also found that UD Epoxy Carbon (230 GPa) Prepreg will have less deformation, as compared to other composites.\nReferences\n[1] P.Y.T. Abkov, E.B. Summers, Lay-up optimization of multilayered anisotropic cylinders based on 3-D elasticity solution, Comput. Struct. 84 (2006) 374\u2013384.\nPlease cite this article as: K. Rajendra Prasad and C. Syamsundar, Theoretical applications, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.201\n[2] Lei Zu, Sotiris Koussios, Adriaan Beukers, Design of filament-wound isotensoid pressure cylinders with unequal polar opening, J. Compos. Struct. 92 (2009) 2307\u20132313. [3] Shafique M.A. Khan, Stress distributions in a horizontal pressure cylinder and the saddle supports, Int. J. Pressure Cylinders Piping 87 (2008) 239\u2013244. [4] A. Alibeigloo, Static and vibration analysis of axi-symmetric angle-ply laminated cylindrical shell using state space differential quadrature method, Int. J. Pressure Cylinders Piping 86 (2008) 738\u2013747. [5] A. Hocine, D. Chapelle, M.L. Boubakar, Experimental and analytical investigation of cylindrical parts of a metallic cylinder reinforced by filament winding while submitted to internal pressure, Int. J. Pressure Cylinders Piping 86 (2009) 649\u2013 655. [6] K.V.J. Raoand, K. Narayana Rao, Design and analysis of filament wound composite pressure cylinder with integrated end domes, Defence Sci. J. 59 (1) (2009) 73\u201381. [7] M.W. Hyer, Stress Analysis of Fiber Reinforced Composite Materials, The McGraw-Hill Companies Inc., USA, 1998.\nand FE analysis of epoxy composite pressure cylinder used for aerospace 9.08.229", + "Further reading\n[8] F.H. Abdulla, M.H. Megat, M.S. Sapuan, B.B. Sahari, Experimental characterization of filament wound glass/epoxy composite materials, ARPN J. Eng. Appl. Sci. 3 (4) (2008). [9] J.C. Velosa, J.P. Nunes, P.J. Antunes, Development of new generation of filament wound composite pressure cylinder, Ciencia e technologia dos Mater. 19(1/2), 2007.\nPlease cite this article as: K. Rajendra Prasad and C. Syamsundar, Theoretical applications, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.201\n[10] Walter H. Tam, Ian A. Ballinger, Jerry Kuo, William D. Lay, Design and Manufacture of a Composite Over Wrapped Xenon Conical Pressure Cylinder, 32ndAIAA/ASME/SAE/ASEE, joint propulsion conference, July 1\u20133,(1996)/lake Buena vista, FL. [11] Thiokol Propulsion, High-pressure Conformable Hydrogen Storage for Fuel Cell Vehicles, proceeding of the (2000) hydrogen review. [12] Bradley S. Forsyth, Christopher Keddy, Honeywell Technology Solutions, Inc., Harold D. Beeson, Filament Winding Simulation of a Composite over Wrapped Pressure Cylinder, prepared for the sample (2001) symposium.\nand FE analysis of epoxy composite pressure cylinder used for aerospace 9.08.229" + ] + }, + { + "image_filename": "designv11_80_0003150_csei50228.2020.9142512-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003150_csei50228.2020.9142512-Figure8-1.png", + "caption": "Figure 8. The interface of the truck crane experiment", + "texts": [], + "surrounding_texts": [ + "When students take a mechanical design course, they can use the virtual remote lab to view the corresponding equipment for course preparation or review, and a step by step from easy to a complex scheme (See Figure5.) is used in building the online experiments. Some experiments interfaces are listed to better illustrate this scheme. For example, when learning drive mechanisms, the virtual lab enables students to operate fundamental experiment at the beginning, after finishing a simple one, a more complex experiment will come, this easy to difficult (see Figure6.) step by step experiments scheme in each individual chapter can consolidate students\u2019 comprehension systematically. a) The interface of the helical gear experiment Also, this step by step scheme is adopted in the whole teaching content. After the learning of each chapter, learners will have a comprehensive and global understanding of the mechanical design course, and to better strengthen the knowledge point and connect the theory of each chapter to practice, several engineering applications are also added at the final part of the course. The below pictures show the experiment interface of a worm screw lift and a truck crane, which includes worm gear and belt drive, coupling and commutator, hydraulic transmission system, the model of internal combustion engines, etc. Students can learn the characteristics of various mechanisms and have a comprehensive understanding of the principles and mechanisms through this application. 250 Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on August 17,2020 at 21:00:29 UTC from IEEE Xplore. Restrictions apply. C. Interavtive opearations In addition, interactive operations are applied to all the experiments, students can drag the left mouse button to rotate the view and drag the right mouse button to pan the view and hold the middle mouse button can zoom the whole view. Through the observation and analysis of these models, the working process of basic mechanisms and equipment becomes more interesting and vivid, which is beneficial for students to form a deeper understanding of the related mechanisms from overall angles. By combining the animations and models in the experiments, mechanical design is no longer a boring and tedious course, rather, 3D animation and interactive operation will attract students to learn and think actively. IV. CONCLUSION In this paper, the teaching process of the online mechanical design course is optimized through the systematic planning of the whole knowledge framework, and corresponding online experiments are designed. This virtual remote laboratory provides a platform where students can get complete training in the whole process of hands-on operation, mechanism and experimental result analysis, thus, students\u2019 awareness of autonomous learning and the ability to simplify and analyze problems will be cultivated. This online experiment is available for students at anywhere and anytime, thus, the contradiction between too many contents and inadequate teaching and experiment time can be solved all at once, in addition, through the interactive operation with mechanisms, mechanical design course will become more interesting and vivid. REFERENCES [1] S. Iqbal, X. Z. Zang, Y. H. Zhu, D. Hussain, J. Zhao, M. M. Gulzar and S. Rasheed, Towards moocs and their role in engineering education, 2015 7th International Conference on Information Technology in Medicine and Education (Itme) (2015), 705-709. [2] Y. Long, M. Zhang and W. F. Qiao, Survey and analysis of the application of massive open online courses (moocs) in the engineering education in china based on a survey of xuetangx, the world's largest mooc platform in the chinese language, Lect Note Netw Syst 22 (2018), 840-850. [3] 3. H. Zhixiu, S. Yongsheng, T. Xiaoying, W. Xiaojing, W. S. O. M. Chengyong, U. Tsinghua, Beijing and China, \"Research to actualize 3d animation on web-based distance learning course in engineering, 2004. [4] G. AlRegib, M. H. Hayes, E. Moore and D. B. Williams, Technology and tools to enhance distributed engineering education, P Ieee 96 (2008), no. 6, 951-969. [5] A. Alexiou, C. Bouras and E. Giannaka, Virtual laboratories in education - a cheap way for schools to obtain laboratories for all courses, by using the computer laboratory, Technology Enhanced Learning 171 (2005), 19-28. [6] R. Vuthaluru, E. Lindsay, N. Maynard, G. Ingram, M. Tade, M. Ang and H. Vuthaluru, Use of digital technologies in bridging the gap between face-to-face and remote engineering programs, 2013 10th International Conference on Remote Engineering and Virtual Instrumentation (Rev) (2013). [7] S. M. Rakshit, S. Banerjee, M. Hempel and H. Sharif, Fusion of vr and teleoperation for innovative near-presence laboratory experience in engineering education, Int Conf Electro Inf (2017), 376-381. [8] V. Potkonjak, M. Gardner, V. Callaghan, P. Mattila, C. Guetl, V. M. Petrovic and K. Jovanovic, Virtual laboratories for education in science, technology, and engineering: A review, Comput Educ 95 (2016), 309-327. 251 Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on August 17,2020 at 21:00:29 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_80_0002407_s43236-020-00062-2-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002407_s43236-020-00062-2-Figure1-1.png", + "caption": "Fig. 1 Configuration of the dual-winding FTPMM: a structure of the stator and rotor; b phasor diagram of the winding arrangement", + "texts": [ + " The phases are physically isolated to prevent the propagation of faults into neighboring phases and to increase thermal isolation. They are also electrically isolated to prevent phase to phase short-circuits and to reduce inverter faults. The windings are magnetically uncoupled to avoid reduced performance in the case of failures of the other phases. A high winding inductance is applied to limit the winding short-circuit current to the rated current [4]. The configurations of the dual-winding FTPMM are shown in Fig.\u00a01a, b. The 12 slots stator consisting of two sets of the identical three-phase concentrated windings are arranged on alternate teeth, namely the ABC and XYZ windings (Fig.\u00a01a). The two sets of the windings of the dual-winding FTPMM used in this study are structured without spatial shifting, and there are no neutral point connections between the winding sets. Windings A and X, B and Y, and C and Z are in the same phase, and the phase difference between the three phases is 120\u00b0 (Fig.\u00a01b). The rotor has eight surface-mounted permanent magnets. To obtain the two-phase d\u2013q-axes electromagnetic parameters 1 3 for all three phases, a Park\u2019s transformation matrix is used in this study as follows: The voltage vectors of the d\u2013q-axes in the synchronous rotating reference frame can be described as follows: where ud and uq are the d\u2013q-axes stator voltages; id and iq are the d\u2013q-axes stator currents; \u03a8d and \u03a8q are the d\u2013q-axes stator flux; \u03a8f is the PM flux; Ld and Lq are the d\u2013q-axes stator inductances; Rs is the resistance of the stator; \u03c9e is the electrical angular velocity of the rotor; and subscript 1 and 2 represent the variables of the ABC and XYZ winding sets, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000569_978-3-030-20751-9_33-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000569_978-3-030-20751-9_33-Figure6-1.png", + "caption": "Fig. 6: Dual interface teleoperation devices", + "texts": [ + " Finally, the velocity feedback v of the CDPR is reflected on both master devices through the admittance controller, and provides the perceptions of the boundary. l\u0307 = L(q)vc (10) In this section, the results of the simulation on the Portable Cable-Driven Robot (PoCaBot) are presented to evaluate the feasibility of the proposed dual interface to achieve accurate and time-efficient teleoperation. A set of experiment results will be presented to show the advantages of the additional trackball. 5.1 Experimental Setup PoCaBot is a spatial cable-driven robot actuated by 8 cables, as shown in Fig. 6a. The experiments and analysis of the PoCaBot were conducted on the simulation platform, Cable-robot Analysis and Simulation Platform for Research (CASPR) [22]. The dual interface includes the customized P2P configured trackball and the modified Novint Falcon to emulate the P2V configured joystick. The way to operate the dual interface is shown in Fig. 6b. Two views shown in Fig. 7 are provided for the user to see and know where the robot is. In the broad view, it is preferable to use the joystick to roughly control the moving direction of the robot. The accurate position control can be achieved for the user by observing the position of the robot in the narrow view. In practice, the range of the magnitude of vd is dominated by the velocity of the P2V joystick vm. The maximum velocity of the joystick is limited. In the experiments, the user was instructed to control the end-effector of the PoCaBot to reach a target position on the xy-plane in three control methods: (1) P2V configured joystick with 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000477_978-3-030-20131-9_72-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000477_978-3-030-20131-9_72-Figure2-1.png", + "caption": "Fig. 2. The US probe guiding module", + "texts": [ + " From structural point of view both modules can be considered as class 3 joints (with 3 DoF) that achieve, at the level of the mobile platform between the Cardan joints a 5 DoF motion ( 0 ). The number of the mobile elements is 3N (the output elements of the two modules as well as the mobile platform). The mechanism has two class 4 joints (the two cardan joints) and two of class 3 joints (the two modules). The number of degrees of freedom is calculated as follows [11]: 1 4 3 5 6 5 3 2 5i i M F N i F C M N C C (1) The second module, for the positioning of the ultrasound probe, shown in Figure 2, has a similar configuration as the first one, the two modules being assembled in a mirrored configuration on a fixed frame to enable the manipulation of both instruments in the same time (figure 3). Each of the two modules of the PRoHep-LCT robotic system has five degrees of freedom with 5 independent parameters for the mobile platform: , , , ,E E EX Y Z . For the inverse kinematic model the coordinates of the characteristic point E, , , , ,E E EX Y Z , are considered known along with all the geometric parameters of the structure, while the values of the active coordinates must be determined: 1 2 3 4 5, , , , " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure45.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure45.1-1.png", + "caption": "Fig. 45.1 Schematic of a circular saw and guide pad b the interaction between the guide pad and circular saw and c distribution of discrete spring-damper elements over the guide pad", + "texts": [ + " Numerically verified models are then subjected to convergence analysis in Sect. 45.4, which includes discussions on the influence of damping on critical speeds, as well as discussion about the influence of modelling discrete spring-damper elements distributed radially, and circumferentially. The paper is concluded in Sect. 45.5. A circular saw, laterally constrained by a guide pad, is modelled as a flexible disc with a clamped inner radius, ri, outer radius, r0, rotating with an angular velocity of X in the clockwise direction as shown in Fig. 45.1. Ignoring any external load, the governing equation of motion of the circular saw in the stationary frame of reference with a guide pad modelled as multiple spring-damper systems, distributed equally over the area of guide pad, is modified from [5] to become: 45 Influence of Guides on Critical Speeds of Circular Saws 521 Dr4u\u00fe qh @2u @t2 \u00fe 2X @2u @t@c \u00feX2 @ 2u @c2 h 1 r @ @r rrr @u @r \u00fe 1 r2 @ @c rc @u @c \u00fe XJ j\u00bc1 kju 1 r d r rj d c cj \u00fe XJ j\u00bc1 cj @u @t 1 r d r rj d c cj \u00bc 0 \u00f045:1\u00de wherein the first term represents bending stiffness, the second term represents inertial stress, the third term represents in-plane rotational stress, and the fourth and fifth terms represent stiffness and damping due to the interaction between the guide and circular saw, respectively, at the location of rj; cj" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001455_012165-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001455_012165-Figure2-1.png", + "caption": "Figure 2. 1st Principal stress ( Max: 35.8 MPa)", + "texts": [], + "surrounding_texts": [ + "IOP Conf. Series: Earth and Environmental Science 343 (2019) 012165 IOP Publishing doi:10.1088/1755-1315/343/1/012165" + ] + }, + { + "image_filename": "designv11_80_0002410_cdc40024.2019.9029501-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002410_cdc40024.2019.9029501-Figure4-1.png", + "caption": "Fig. 4. Fixed camera configuration for robot visual servoing.", + "texts": [ + "352, the closed-loop system becomes unstable. This suggests also that sharper methods for determining sign definiteness of V\u0307 could be sought to determine less conservative conditions for the controller gains and to allow a larger upper bound \u00b5 for the norm of Kpa the anti-symmetric part of Kp. The benchmark example can be brought to a practical scenario of visual servoing problem. Consider a planar manipulator monitored by a fixed camera with optical axis orthogonal to the robot workspace plane as in [21], [22] (see Figure 4). If the manipulator is assumed to have negligible dynamics, i.e., a kinematic manipulator, then the mathematical control problem boils down to the control of a multivariable two-dimensional integrator of the form: \u03c3\u0307 = B\u03bd, \u03bd = Spu, B = R(\u03b8) (17) where B represents the rotation matrix relating the camera frame to the manipulator workspace frame both illustrated in Fig. 4. The possible scaling factors between the two frames are assumed unitary for simplicity. The rotation matrix R(\u03b8) is given by: R(\u03b8) = [ cos(\u03b8) sin(\u03b8) \u2212 sin(\u03b8) cos(\u03b8) ] (18) For the uncalibrated camera case, B = R(\u03b8) is not exactly known except for some nominal value of \u03b8, say \u03b8n. Let \u2206\u03b8 = \u03b8\u2212\u03b8n. Then, following the proposed method, Sp = R(\u2212\u03b8n) is chosen so that Kp = BSp = R(\u03b8)R(\u2212\u03b8n) = R(\u2206\u03b8). Now, Kp = R(\u2206\u03b8) can be decomposed as done in the benchmark example, Kp = cos(\u2206\u03b8)I + sin(\u2206\u03b8) [ 0 1 \u22121 0 ] (19) Here, it is easy to see that \u00b5 \u2265 sin (\u2206\u03b8), and \u03bbm = \u03bbM = cos (\u2206\u03b8)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002696_s40430-020-02368-5-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002696_s40430-020-02368-5-Figure15-1.png", + "caption": "Fig. 15 Transmission system of cutting gearbox", + "texts": [ + " The source of vibration excitation might be internal excitation caused by gear meshing and external excitation caused by coal cutting. According to the literature [23], the frequency of coal cutting load is about two times the rotation frequency of the cutting drum. In this paper, the rotation frequency of drum for MG500-1180/WD coal mining machine is 0.5\u00a0Hz. Therefore, the frequency of external excitation is about 1\u00a0Hz, and it is impossible for external excitation to cause resonance of CPGH. Resonance is caused by internal excitation, and the gear meshing frequencies will be calculated. Figure\u00a015 shows the transmission system of cutting gearbox. The system consists of two-stage straight tooth gear 1 3 transmission, two-stage planetary gear transmission, motor and cutting drum. Based on transmission system parameters provided by factory, the meshing frequencies of straight tooth gear transmission are: where n and z are rotation speed of gear axis (r/min) and the number of teeth, respectively. (21)fz = nz\u221560 1 3 The meshing frequencies of planetary gear transmission are [24]: where zc , za , fb , fa are teeth number of ring gear, teeth number of sun gear, rotation frequency of carrier, rotation frequency of sun gear axis, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure80.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure80.3-1.png", + "caption": "Fig. 80.3 Application of boundary conditions and load parameters. Two rigid plates, red and blue are fixed at top and bottom of designed scaffold structure and compressive load of 75 N is applied in negative Y direction", + "texts": [ + " The elastic isotropic material properties of Ti6Al4V adopted in the study include Young\u2019s modulus of 114 GPa, Poisson\u2019s ratio of 0.33, density of 4.43 g/cc, and compressive yield strength of 1070 MPa [19]. Remeshing with 4-node SHELL181 hexahedral element was performed for each model (Fig. 80.2). Two rigid plates are fixed at the top and bottom of each scaffold to assign the boundary conditions and loading. The bottom plate is constrained in all directions while a static compressive load of 75 N is applied in negative Y direction at the top plate (Fig. 80.3). 80 Triply Periodic Minimal Surface-Based Porous Scaffold Design \u2026 959 960 V. Rati et al. Porosity was calculated from the relation expressed in Eq. (80.3) [20]. N \u00bc 1 VP VS \u00f03\u00de where VS is the solid volume and VP is the volume of porous construct. The solid volume constitutes the total volume of cubic construct. Solid volume (VS) and porous volume (VP) were calculated all primitive (P1\u2013P7) and gyroid surface (G1\u2013 G7) models with different porosity levels (60\u201390%) as listed in Table 80.2. The effective elastic modulus as well as compressive strength was calculated for each model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003614_j.procir.2020.04.092-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003614_j.procir.2020.04.092-Figure8-1.png", + "caption": "Fig. 8. The discrete level set representations of the used part (\u03a6\ud835\udc62\ud835\udc62) and final part (\u03a6\ud835\udc53\ud835\udc53).", + "texts": [ + " \u03a6\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305 = (\u03a6\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 + t) \u2216 \u03a6\u0303i (22) \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 = \u03a6\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305 \u2216 \u03a6\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 (23) \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 \u2192 \ud835\udc35\ud835\udc35\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 (24) \ud835\udc35\ud835\udc35\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 \u2192 {SF\ud835\udc5b\ud835\udc5b+1, SF\ud835\udc5b\ud835\udc5b+2, SF\ud835\udc5b\ud835\udc5b+3, . . . SF\ud835\udc5b\ud835\udc5b+\ud835\udc5a\ud835\udc5a} (25) In this section, the proposed method is validated by a remanufacturing case study. There is a used part (Fig. 7a) required to remanufactured to a final part (Fig. 7b). The CAD models of the used part and final parts are modelled by discrete level set functions as \u03a6\ud835\udc62\ud835\udc62 and \u03a6\ud835\udc53\ud835\udc53 , separately. Both of them are built on a design domain D (150*150*100) with grid size \u2206\ud835\udc65\ud835\udc65 (0.5 mm), as shown in Fig. 8. The next step is the maximization of the intersection volume, so the intersection part extraction algorithm is implemented for the used part. As Fig. 9 shown, the used part is transformed from \u03a6\ud835\udc62\ud835\udc62 to \u03a6\u0303\ud835\udc62\ud835\udc62 by the optimal transformation matrix obtained by the proposed algorithm and then the intersection part \u03a6\ud835\udc56\ud835\udc56 is obtained. The result of the optimal transformation is shown in Table 3. After obtaining the intersection part, the next step is modifying the intersection part from \u03a6\ud835\udc56\ud835\udc56 to \u03a6\u0303\ud835\udc56\ud835\udc56 by respecting different AM processes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002375_s11768-020-9125-2-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002375_s11768-020-9125-2-Figure4-1.png", + "caption": "Fig. 4 Schematic diagram of a two-link manipulator.", + "texts": [ + " The stable range of the gain increases as the size of sample time decreases under the same reference; in addition, as the frequency of the sinusoidal trajectory increases, the stable gain range decreases. These results are consistent with observed experimental results in actual TDC systems. under the 5th order polynomial trajectory, equation (54). (b) T = 0.005 s under the trajectory sin( 1 2 t), equation (55). (c) T = 0.010 s under the trajectory sin(2t), equation (56). In this section, dynamic simulations for a two degreesof-freedom manipulator are performed to validate the proposed stability analysis of the standard discrete TDC as an example of a practical nonlinear system application. Fig. 4 illustrates the two-link manipulator described by the following dynamics: \u03c4 =M(\u03b8)\u03b8\u0308 + C(\u03b8\u0307,\u03b8) + g(\u03b8), (57) where \u03c4 \u2208 R2 denotes the joint torque vector, \u03b8 \u2208 R2 denotes the joint angle vector, M(\u03b8) denotes the manipulator inertia matrix, C(\u03b8\u0307,\u03b8) denotes the centrifugal and Coriolis torque matrix, and g(\u03b8) denotes the gravitational torque vector, while their elements are given as M11 = m1l2c1 +m2(l21 + l2c2 + 2l1lc2 cos\u03b82) + I1 + I2, M12 =M21 = m2(l2c2 + l1lc2 cos\u03b82) + I2, M22 = m2l2c2 + I2, C1 = \u2212m2l1lc2 sin\u03b82(2\u03b8\u03071\u03b8\u03072 + \u03b8\u0307 2 1), C2 = m2l1lc2 sin\u03b82\u03b8\u0307 2 1, g1 = {(m1lc1 +m2l1) cos\u03b81 +m2lc2 cos (\u03b81 + \u03b82)}g, g2 = m2lc2 cos (\u03b81 + \u03b82)g, where (m1, l1, lc1, I1), (m2, l2, lc2, I2) denote the mass, link lengths, position of the centre of mass, and the moment of inertia of links 1 and 2, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000280_s10999-019-09455-z-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000280_s10999-019-09455-z-Figure9-1.png", + "caption": "Fig. 9 Modified epitrochoid generated starring", + "texts": [ + ", R0 R0 r0\u00f0 \u00de \u00bc Z \u00f013\u00de The coordinates of point M (Xm, Ym) of the modified epitrochoid in dimensionless form, with respect to XOY co-ordinate system, are as follows: Xm \u00bc A0 cosw\u00fe 1 Z cos Zw rm cos w\u00fe /\u00f0 \u00de Ym \u00bc A0 sinw\u00fe 1 Z sin Zw rm sin w\u00fe /\u00f0 \u00de \u00f014\u00de The leaning angle (/) is an important parameter in the modification of epitrochoid profiles because in gearing action (p=2 /) may be considered as the instantaneous pressure angle. In dimensionless parametric form it is expressed as: / = tan 1 sin Z 1\u00f0 \u00dew A0 \u00fe cos Z 1\u00f0 \u00dew \u00f015\u00de where w \u00bc Zn \u00f016\u00de Figure 9 shows a typical epitrochoid generated star-ring (with roller) set. Appendix 2 Calculation of side thrust on the ring due to tightening it between the end plate and valve plate by fixing bolts (Fig. 1c) for Type-C FEM model Each side flat area of the ring of this considered ORBIT motor unit is 3143.17 mm2, as estimated by solid modelling. Each bolt (M8 9 1.25) cross sectional (effective) area As, is 36.6 mm2. Number of bolts 7. We assume the Proof Strength (ry) as 1098.34 MPa considering the steel\u2019s strength class 12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001934_978-981-15-1124-0_17-Figure21-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001934_978-981-15-1124-0_17-Figure21-1.png", + "caption": "Fig. 21 Mode Shape 7", + "texts": [], + "surrounding_texts": [ + "shown in images Fig. 15 and Fig. 24. We can notice that the fracture of the femur bone seems to occur either at the shaft or the neck region for different frequencies. Although an external load can excite any mode from the first mode to any higher mode, the lower nodes are comparatively easier to excite." + ] + }, + { + "image_filename": "designv11_80_0002756_012052-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002756_012052-Figure1-1.png", + "caption": "Figure 1. Cross-sectional view of the beam. Figure 2. The vehicle model.", + "texts": [ + " The model of the passenger car is established in ADAMS/Car, and is optimized based on the mathematical model of the optimization and related theories. The front suspension of the car studied in this paper is the McPherson independent front suspension, and the rear one is the torsion beam semiindependent rear suspension. The main parameters of the vehicle are as follows: The torsion beam suspension model of the flexible body structure is established according to the given parameters in order to obtain the ride comfort data on the driver's seat, and then the finite element model shown in figure 1 is established. The finite element mesh of the torsion beam suspension is divided, then constrained according to the condition proposed above. And finally assigning parameters to the constraints and performing analytical calculations. 6th International Conference on Advanced Engineering and Technology (ICAET 2019) IOP Conf. Series: Materials Science and Engineering 811 (2020) 012052 IOP Publishing doi:10.1088/1757-899X/811/1/012052 Performing modal analysis in HyperMesh, calculate the modal neutral file (" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001881_s12206-019-1113-4-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001881_s12206-019-1113-4-Figure9-1.png", + "caption": "Fig. 9. Measurement positions of residual stress (CASE-TS).", + "texts": [ + " We measured the residual stress on the critical cross sections of the teeth faces using an X-ray residual stress analyzer (PULSTEC \u03bc-X360s). As the residual stress in CASE-TS, we measured the residual stress of a supporting gear for each helix angle. We measured the residual stress of four teeth of each target gear. We measured the residual stress at 13 measurement points: 0.4 mm, 2.8 mm, 5.2 mm, 7.6 mm, 10.0 mm, 12.4 mm, 15.0 mm, 17.6 mm, 20.0 mm, 22.4 mm, 24.8 mm, 27.2 mm, and 29.6 mm from ACUTE-END in the tooth width direction for each tooth, as shown in Fig. 9. Hence, we measured a total of 52 residual stresses for each helix angle in CASE-TS. As the residual stress in CASE-T, we measured the residual stress of a test gear for each helix angle. We measured the residual stress of four teeth of each target gear. We measured the residual stress at nine measurement points: 0.4 mm, 2.8 mm, 5.2 mm, 7.6 mm, 10.0 mm, 12.4 mm, 14.8 mm, 17.2 mm, and 19.6 mm from ACUTE-END in the tooth width direction for each tooth. Hence, a total of 36 residual stresses for each helix angle in CASE-T were measured" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002512_icase48783.2019.9059099-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002512_icase48783.2019.9059099-Figure4-1.png", + "caption": "Fig 4: Shows the Arrangement made by Goff for wind tunnel testing of Brazuca [3]", + "texts": [ + " The experiments were conducted at University of Tsukuba. Maximum velocity of the wind that could be achieved in the tunnel was 55 m/s. The outlet of blower was 1.5 meters \u00d7 1.5 meters and turbulence was less than 0.1%. Diameter of the soccer balls that were tested was 0.22 meters, which translates to the blockage of around 1.7 %. He used a conventional rear-mounted method [10], in which air flows over the ball in a direction opposite to the mounting of horizontal stainless steel rod. This set-up is shown in figure 4 where a Brazuca soccer ball is under experimentation: Goff et al [3] performed experiments for two orientations for each selected soccer ball. His selection for orientation was based on significant difference in panels on frontal geometry to flowing air. Aerodynamic parameters were measured at flow velocities of 7-35 m/s which translates to Reynolds numbers of 105-525, where Re = vD/\u03bd [11], with D the ball\u2019s diameter, and \u03bd = 1.54 \u00d7 10\u22125 m2/s, the kinematic viscosity. For a nine second interval, forces on the soccer ball under test were measured using a six component force balance with sting (model number LMC-61256 by Nissho Electric Works Co, Ltd)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001197_s11249-019-1224-1-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001197_s11249-019-1224-1-Figure4-1.png", + "caption": "Fig. 4 a Shear rate distributions at different positions of the slider. b The film thickness along the direction of the rotating disk, the pseudo interference fringe indicating the same film thickness, and the yellow line indicating the rotating direction, slider size (Breadth \u00d7 Length) 4 \u00d7 6\u00a0mm, l = 27\u00a0mm (Color figure online)", + "texts": [ + " The inclination angle between the slider and the transparent disk is \u03b8, and disk rotation speed is \u03c9. When the slider is in the center of the disk, the shear strain rate from the center r is equal to \u22c5 = \u03c9 \u00d7 r/ (tan(\u03b8) \u00d7 r) \u2248 \u03c9/\u03b8, which means that shear rate is approximately constant. It can be seen that the shear rate tends to be constant at a large film thickness. When the distance from the center point of the disk to the slider is l, the shear rate at a distance r from the center of the disk is given by \u22c5 = \u03c9 \u00d7 r/(tan(\u03b8) \u00d7 (r \u2212 l)) \u2248 \u03c9/\u03b8 \u00d7 (1 + l/(r \u2212 l)), as shown in Fig.\u00a04a. The observed photobleached area is relatively small compared to the dimensions of the slider. The shear rate can be assumed to be locally invariant in the observed photobleached area. The path of the rotatory disk is a curve, and equal film thickness position is a straight line, as presented in Fig.\u00a04b. There is a velocity component in vertical direction of the slider\u2013disk contact line. Divergent velocity component v1 will not create hydrodynamic lubrication, while the velocity component v2 along the convergent gap direction will generate slight hydrodynamic pressure. However, the velocity component is so small as to be negligible. In experimental condition, velocity profiles are the almost same along the equal film thickness direction, which is verified by the experimental results, regardless of whether the velocity component is convergent or divergent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.6-1.png", + "caption": "Figure 1.6. Illustration of the formula v\u0307r = ar ! !r \" vr when v\u0307r = 0. Left: the robot follows a circle; right: representation of the vectors in the robot frame", + "texts": [ + "R1/R0|R1 = ($x,$y,$z) corresponding to the rotation vector of the box relative to R0 expressed in R1. We have to express a, ! in the coordinate system of the box since these quantities are generally measured via the sensors attached on it. The first state equation is: p\u0307 [1.8] = R (&, !,') \u00b7 vr. Let us differentiate this equation. We obtain: p\u0308 = R\u0307 \u00b7 vr +R \u00b7 v\u0307r with R = R (&, !,'). From this equation, we isolate v\u0307r to get v\u0307r = RT \u00b7 p\u03082 34 5 ar \"RTR\u0307 \u00b7 vr [1.4] = ar \" !r ' vr, which constitutes the second state equation. Figure 1.6 shows a situation where a robot has a constant speed and follows a circle. The speed vector vr is a constant whereas ar is different from zero. We also need to express - &\u0307, !\u0307, '\u0307 . in function of the state variables. From equation [1.11], the relation !|R0 = R (&, !,') \u00b7 !|R1 becomes $ % cos ! cos' \" sin' 0 cos ! sin' cos' 0 \" sin ! 0 1 & ' $ % &\u0307 !\u0307 '\u0307 & '=R (&, !,') \u00b7 !r. By isolating the vector - &\u0307, !\u0307, '\u0307 . in this expression, we obtain the third state equation. By bringing together the three state equations, we obtain the navigation mechanization equations [FAR 08]: ( 9999) 9999* p\u0307 = R (&, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure8-1.png", + "caption": "Figure 8 Second mode.", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0000530_978-3-030-12082-5_55-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000530_978-3-030-12082-5_55-Figure8-1.png", + "caption": "Fig. 8 Distribution of temperature on contact surface of platform, \u00b0C", + "texts": [], + "surrounding_texts": [ + "A number of computational experiments were carried out, the conditions and results of which are summarized in Table 2. The temperature field of the macromodel was determined, and then, to compare experiment results, the thermal contact conductance of the gyro unit-platform contact region averaged over the nominal area was calculated by the relation \u03b1c q\u0304 T2 \u2212 T1 where q\u0304 is the average heat flux density over the area of finite elements of the lower nominal surface, T1, T2 are the average temperatures over the area of finite elements of the nominal lower and upper contact surfaces, respectively. Experiments were conducted to evaluate the influence of various factors, while the most appropriate model should be considered as the experiment No. 4 model. The TCC parameter was set as a constant value equal to 157,000 W/(m2 K) or as the above-tabulated dependence on pressure TCC(p). The value 157,000 W/(m2 K) is obtained from the graph in Fig. 5 for the average pressure from the bolt clamp force of 7.3 MPa, calculated by dividing the sum of the clamp forces of each bolt (2000 N) by the nominal contact area. To evaluate the effect of thermal expansion, experiments were carried out for two types of contact behavior, Standard and No separation. For the Standard contact type, the contact heat transfer occurred strictly in the real contact area (Fig. 6), for which the TCC parameter was set. Thus, this type of contact reflects the influence of change in shape from thermal expansion. For the No separation contact type, movement of the contact surfaces along the contact plane is allowed, but separation of the surfaces is not permitted and the real contact area is equal to the nominal one. Thus, in the case of No separation contact, the change in shape of the surfaces from thermal expansion is not reflected in temperature results since it does not affect the thermal contact conductance. In this case, the heat transfer occurs over the entire nominal contact area. The wide use of this type of contact in actual practice is due to the significantly lower computational complexity and, accordingly, solution time. The No separation contact type was set on the gyro unit-platform connection in experiments Nos. 1 and 2. The calculations were carried out with the assumption of small displacements, since the accounting for large displacements for experiment No. 6 resulted in a change in the averaged thermal contact conductance of 0.1%, which is considered insignificant. The first and second experiments set the heat transfer throughout the whole of the nominal contact area. In this case, setting the dependence of the thermal contact conductance TCC on the contact pressure p obtained in the micromodel led to a decrease of 4.3 times in the averaged thermal contact conductance of themacromodel. The models used in the computational experiments Nos. 3 and 4 take into account the effect of thermal expansion on the real contact area. Because of the change in shape of the cylindrical body of the gyro unit, tangency takes place in the form of a narrow ring along the outer edge of the nominal contact area.Also, areas near the bolts are in direct contact. The real contact area was 56% of the nominal area (Fig. 6). As is clear from a comparison of experiments Nos. 3 and 1, the averaged thermal contact conductance decreased bymore than 5 times just due to accounting for the real contact area at a constant TCC of 157,000 W/(m2 K). Under the same conditions and using the TCC(p) dependence (experiments Nos. 2 and 4), the averaged thermal contact conductance decreased noticeably less, by 56%, which can be considered a result of thermal expansion without the direct influence of contact pressure. Repetition of the result of 56% is a random coincidence in this case. Experiment No. 5 showed that the use of a friction coefficient 0.3 instead of 0.5 led to a slight increase in the thermal contact conductance (by 18%). Thus, the friction coefficient has a noticeable effect on the conductance of the actual contact. The thermal contact conductance is significantly affected by clamp force of the bolts. Experiment No. 6 showed that using the conditions of experiment No. 4 and decreasing the clamp force from 2000 to 100 N, the averaged thermal contact conductance decreased by more than 6 times. A similar effect in real structures can arise in the case of more complicated connections, for example, with clasps [35]. The distribution of contact pressures, surface temperatures, and heat fluxes for the contact platform surface in experiment No. 4 is shown in Figs. 7, 8, and 9. Figure 10 shows the distribution of temperatures throughout the entire model of the gyro unit-platform assembly under the same conditions." + ] + }, + { + "image_filename": "designv11_80_0000303_978-981-13-3627-0-Figure5.24-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000303_978-981-13-3627-0-Figure5.24-1.png", + "caption": "Fig. 5.24 Stress contour of the chip subjected to the impact effect induced by the ejector needle at a speed of 0.1m/s under the through-penetration condition.\u00a9 2011 IEEE.Reprinted, with permission, from Ref. [39]", + "texts": [ + " Local regions surrounding the contact points are refined to ensure the accuracy of numerical results. The dynamic impact force acts directly on the chip in the through-penetration case, resulting in a high instantaneous stress concentration at the local contact area on the backside of the chip. The impact speed is an important factor affecting the stress level. The VonMises stress contour of the chip subjected to the impact force induced by the ejector needle at a speed of 0.1 m/s under the through-penetration condition is shown in Fig. 5.24. It indicates clearly that the most severe stress occurs at the back center of the chip. However, the high stress is only restricted to a very limited region surrounding the contact center, and most of other regions are nearly in stress-free state. Moreover, the stress due to needle impacting can be 8\u201310 times larger than static stress, and the stress decreases remarkably as the distance from the contact center enlarges. For the free-penetration case, a direct contact exists between the ejector needle and the substrate, and the impact effect is transmitted through the substrate to the chip", + " In contrast, if the substrate is through- 134 5 Single-needle Peeling penetrated during the peeling-off process, it may lead to excessive local stress. The substrate penetration should be avoided to prevent chip damages. It should be pointed out that the calculationmethod abovementioned can be used to select suitable impact speed. The variation of the third principle stresses for four elements around the contact center under the through-penetration condition is depicted in Fig. 5.26. The element A is located at the contact center, the elements B, C, and D are arranged gradually away from the element A, as illustrated in Fig. 5.24. It can be seen that the third principle stress of element A decreases steeply from zero to nearly \u221238.6 MPa upon the contact, and then the stress level is reduced obviously. Subsequently, a second stress valley is observed at the element A due to the second impact effect of the ejector needle, despite that the stress level is lower than the first valley. In addition, the impact effect on the elements B, C, and D is insignificant, which can be derived from the relatively small variation of the third principle stress" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003972_icma49215.2020.9233579-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003972_icma49215.2020.9233579-Figure4-1.png", + "caption": "Fig. 4 Mechanism of the spoke.", + "texts": [ + " The foundation is designed with a rail structure that allows the spokes to slide stably as shown in Fig. 3. The picture of the assembled robotic wheel based on the design drawing is shown in Fig. 1. The robotic parts are printed by a 3D printer (AGILISTA-3200, KEYENCE) and also fabricated by a modeling plotter (NC-5SK, Mimaki). The materials of the parts are acrylate resin and ABS. 2) Rack & Pinion Mechanism A pinion gear connected with a servo motor is combined with a rack gear with a spoke as shown in Fig. 4. The extension can be achieved by the rack & pinion mechanism up to the half length of a radius. Each spoke moves in the direction of extension from the initial radius of the wheel and 1258 Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 19,2021 at 09:15:10 UTC from IEEE Xplore. Restrictions apply. returns to the initial length. A pinion gear is attached to the shaft of a servo motor, and the rotating motion is transmitted to a rack gear with the spoke. The rotation of the pinion gear is controlled by a servomotor, and the spoke extends by following clockwise rotation from the origin of rotation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003307_aim43001.2020.9158857-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003307_aim43001.2020.9158857-Figure1-1.png", + "caption": "Fig. 1 Prototype of the linear series elastic actuator", + "texts": [ + "00 \u00a92020 IEEE 960 Authorized licensed use limited to: Carleton University. Downloaded on August 27,2020 at 13:02:57 UTC from IEEE Xplore. Restrictions apply. presents the typical impedance and admittance controller and shows the dependence of the virtual stiffness on the environmental dynamics. The modified impedance and admittance controllers are developed and analyzed in Sec. IV. Finally, conclusions are given in Sec. V. To implement the proposed new impedance and admittance controllers, we consider a linear series elastic actuator as shown in Fig. 1. This SEA consists of a two-phase bipolar stepper motor (Oriental Motor PKP214) that can produce a maximum torque of 0.036 Nm. The output of the stepper motor is connected to a ball screw with a diameter of 4 mm and lead of 1 mm. The rotation of the ball screw is further transmitted to the linear motion of a nut. The nut is then connected to the output bracket through two identical side springs. The output bracket and nut are both constrained by a linear guide. Fig. 1 also shows the side springs. Unlike previous SEAs that used commercially available coils springs [10, 15], the costumed-made side spring has a planar structure that can be easily miniaturized to be adapted for the compact space between the screw nut and the output bracket. The side springs are made of annealed plastic mold steel (S-STAR) and the two side springs together provide a stiffness of 110.38 N/mm in both tension and compression directions. The deformation of the side springs is measured using a linear optical encoder that has a resolution of 1 \u03bcm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001670_icems.2019.8922498-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001670_icems.2019.8922498-Figure1-1.png", + "caption": "Fig. 1. Main Combination Modes of Induction Motor. (a) Mode of Stator Core. (b) Mode of Interference Fit. (c) Mode of bearing contact", + "texts": [ + " So the influence of the modal simulation precision of the induction motor is actually the contact between the components, which is summarized from the transmission path of the vibration as the laminated contact between the core punching pieces, interference contact between the stator core and the casing, and bearing contact between the shaft and the end cap. In this paper, the main modal parameters, including the orthotropic material properties of the stator, the contact stiffness factor of the stator and the casing, the bearing stiffness, are used as the optimization parameters, and the modal of the components of the Y series induction motor are actually tested. In order to avoid interference with each other, univariate combination model is established for modal analysis, as shown in Figure 1. The measured modal frequencies of the corresponding modes are set as the simulation optimization targets, and optimized, so that the key modal parameters suitable for the Y series induction motor are obtained, and the accuracy of the modal parameters is verified by the test results. II. THE KEY PARAMETERS OF MODAL ANALYSIS The stator of the machine is the main path for the vibration of electromagnetic noise. The structure of the stator 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEEAuthorized licensed use limited to: University of Canberra" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003645_icept50128.2020.9203003-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003645_icept50128.2020.9203003-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of laser heating and infrared temperature measurement used in laser jet solder ball bonding and processes.", + "texts": [ + " 2020 21st International Conference on Electronic Packaging Technology (ICEPT) temperature change of the melted solder during the heating process in the nozzle is obtained. In the simulation, the temperature field of the solder ball in the nozzle during the laser jet soldering process was obtained by the transient thermal method. II. EXPERIMENTAL PROCEDURE The diameter of the solder ball used in the laser jet solder ball bonding process experiment was 0.6 mm and the material was Sn3.5Ag0.5Cu (SAC305). First, the solder ball was put into the nozzle, and then the laser is turned on to heat and measure the temperature of the solder ball. Fig. 1 shows the schematic diagram of laser heating and infrared temperature measurement used in laser jet solder ball bonding and process. In the welding experiment, a fiber laser with a total power of 130 W, a wavelength of 1064 nm, a spot size of 300 um, and a minimum detection temperature of infrared thermometer of 453 K are used. The total heating time of the laser is 9 ms, the first 5 ms use 117 W power heating, the last 4 ms use 104 W power heating. Fig. 2 shows the welding equipment used in the experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000532_itoec.2018.8740752-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000532_itoec.2018.8740752-Figure4-1.png", + "caption": "Fig. 4. a Robot positions along Z direction. b Robot positions in the tests", + "texts": [ + " For each direction, three times of the impact tests were repeated acting on the end-effector of the robot. The data acquisition system was completed using the SignalPad system based on LabVIEW. The sampling frequency was set at 5.12 kHz, the frequency measurement band was 0-200 Hz, and the frequency resolution was 1 Hz. In order to investigate the dynamic behavior of the robot in different configurations, the robot was moved on the range of its limits along three directions X, Y, and Z, and totally seven configurations were investigated as shown in Fig. 4. The position of the home position (X2, Y2, Z2) in Fig. 4 was used as the benchmark. As an example, the FRF the home position are plotted in Fig. 5. The data imply that 16 Hz is close to the natural frequency of the robot. Now we compare the dynamic performance index at the frequency of 16 Hz and 40 Hz. As Show the vibration test results in Table I, the dynamics performance index at 16 Hz are significantly larger than those at 40 Hz for all the above seven configurations. This indicates that when the robot is subjected to an external force of 16 Hz, the structural dynamic performance is poor for the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002705_kem.841.144-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002705_kem.841.144-Figure1-1.png", + "caption": "Fig. 1 The spur gear to be investigated", + "texts": [ + " No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications Ltd, www.scientific.net. (#541197814, University of Melbourne, Melbourne, Australia-26/07/20,01:08:57) Material. This study examined the spur gears of an incapacitated hand tractor after operating for only 40 hours, classified as premature failure. This research was conducted using empirical methods and the finite element method (FEM). The subject of this study is illustrated in Fig. 1. Visual Inspection. Initially, a detailed visual inspection of the hand tractor was conducted to assess the general quality of the gears and to identify all relevant fracture features, as displayed in Fig. 2. Chemical Composition Analysis. Chemical analysis was also carried out to determine the chemical composition of the gears so that it could be classified correctly into the proper standard groups, such as ASM or AISI. Hardness Testing. Hardness testing was performed at 9 points on the outer and inner surface of the specimen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure7.18-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure7.18-1.png", + "caption": "Figure 7.18 Schematic of triangular membrane.", + "texts": [ + "1) to the sine transforms and accomplishing all operations in the same way as it was done using the Fourier transformation, we obtain Eq. (7.7.17), in which Green\u2019s function will already take the form of the discrete expansion. G(x, x\u2217; y, y\u2217) = 4 Tlxly \u221e\u2211 m,n=1 sin \ud835\udc5a\ud835\udf0b\ud835\udc65 lx sin \ud835\udc5b\ud835\udf0b\ud835\udc66 ly sin \ud835\udc5a\ud835\udf0bx\u2217 lx sin \ud835\udc5b\ud835\udf0by\u2217 ly (\ud835\udc5a\ud835\udf0b lx ) 2 + (\ud835\udc5b\ud835\udf0b ly ) 2 + \ud835\udf062 (7.7.24) The spectral expansions in the finite domain are also effective during the arrangement of the compensating actions on the border of the extended domain. As an example of the application of SMBE let us examine the solution of Eq. (7.7.1) in the triangular domain, shown in Figure 7.18. Let us assign boundary conditions in the form of (7.7.2), we assume that the function f(x, y) on the sides of triangle x = 0 and y = 0 becomes zero, and the diagonal is given according to the law of the triangle. As the extended domain we assume a rectangle with the sides lx and ly, on which boundary conditions (7.7.23) are satisfied. The Green function will take the form (7.7.24). In this case the solution of equation will be written down in the form of: w(x\u2217, y\u2217) = wq(x\u2217, y\u2217) + 4 \ud835\udc47\ud835\udc59xly \u221e\u2211 m=1 \u221e\u2211 n=1 g\ud835\udc5a\ud835\udc5b(x\u2217, y\u2217) \u00d7\u222b \u0393\u2032 \ud835\udf11(x, y) sin \ud835\udc5a\ud835\udf0b\ud835\udc65 lx sin \ud835\udc5b\ud835\udf0b\ud835\udc66 ly d\u0393\u2032 (7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002435_012137-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002435_012137-Figure5-1.png", + "caption": "Figure 5. Commutator Illustration in DC Motor", + "texts": [ + " And the torque produces by the motor inertia is the inertia multiplied by the rate of angular speed change, as shown in equation 6. \ud835\udc49\ud835\udc59 = \ud835\udc3f \ud835\udc51\ud835\udc56 \ud835\udc51\ud835\udc61 (5) \ud835\udc47\ud835\udc56\ud835\udc5b\ud835\udc52\ud835\udc5f\ud835\udc61\ud835\udc56\ud835\udc4e = \ud835\udc3d\ud835\udc40 \ud835\udc51\ud835\udf14 \ud835\udc51\ud835\udc61 (6) In Figure 4, we added the inductance to the DC Motor model. The 3rd International Conference on Eco Engineering Development IOP Conf. Series: Earth and Environmental Science 426 (2020) 012137 IOP Publishing doi:10.1088/1755-1315/426/1/012137 If we look closely into the design of commutator, it actually switches on and off the voltage source, as the rotor coils rotates through the commutator (as illustrated in Figure 5). Figure 6 shows the circuit that does the modelling of the commutator. The 3rd International Conference on Eco Engineering Development IOP Conf. Series: Earth and Environmental Science 426 (2020) 012137 IOP Publishing doi:10.1088/1755-1315/426/1/012137 Here, we discuss the results of modeling DC Motor in LTSpice. The results are divided into 3 gradual steps, to foster better understanding of how a DC Motor works. By using the circuit modelling of Figure 2, we can see that the simulation output for speed and backemf voltage are exactly following the voltage source, as shown in Figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002696_s40430-020-02368-5-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002696_s40430-020-02368-5-Figure2-1.png", + "caption": "Fig. 2 Force analysis of gearbox housing", + "texts": [ + " Coal mining is forced to be in deep and complex geological condition coal seams with the depletion of the shallow coal resources, and deep mining will cause inconvenience in equipment maintenance. Therefore, the reliability of coal mining machine is one of the most concerned problems of coal mining companies. However, various unexpected problems have been observed during the cutting operation, and CPGH is one of the weak parts [1]. Recent problems [2, 3] involve the fracture failure of housing during operation in section A\u2013A and B\u2013B (Fig.\u00a02), resulting in reduction in coal production. CPGHs are different from traditional housings (automotive, aerospace, industrial, wind turbine gearboxes, et\u00a0al.) which are designed to be like cantilever beams for the objective of coal cutting as shown in Fig.\u00a03. A CPGH consists of motor shell, middle part of gearbox housing and planetary gear system shell (Fig.\u00a02). The excitations of CPGHs involve not only high-frequency internal vibrations caused by gear meshing, but also low-frequency heavy external vibrations caused by coal cutting. Stress concentration and excessive deformation can be easily caused by the cantilever beam shape of CPGHs. Previous Technical Editor: Wallace Moreira Bessa, D.Sc. * Yimin Zhang zhangyimin@syuct.edu.cn 1 School of\u00a0Mechanical Engineering and\u00a0Automation, Northeastern University, Shenyang\u00a0110819, China 2 Equipment Reliability Institute, Shenyang University of\u00a0Chemical Technology, Shenyang\u00a0110142, China 3 Shenyang Engine Design Institute of\u00a0China Aero Engine, Shenyang\u00a0110000, China 1 3 Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:273 1 3 Page 3 of 15 273 papers [4\u20137] indicated that fatigue failure is due to the joint action of local resonance and stress concentration and the failure region is often the part of stress concentration and large strain in local resonance", + " Bending moment in Z direction (7)FNA = FNB = Px cos \u2212 ( Pz \u2212 G sin ) (8)MyA = [( Pz \u2212 G cos ) + Px sin ] LA \u2212Mp (9)MyB = [( Pz \u2212 G cos ) + Px sin ] LB \u2212Mp (10) MzA = [ Px cos \u2212 ( Pz \u2212 G ) sin ] L 7 \u2212 A 1 ( L 0 + LA ) (11) MzB = [ Px cos \u2212 ( Pz \u2212 G ) sin ] L 7 \u2212 A 1 ( L 0 + LB ) N 3y Y direction reaction force of hinged joint N 4y Y direction reaction force of hinged joint G Gravity of CPGH L 0 Eccentricity of A1 Pz Vertical force of cutting load A Normal stress of A\u2013A B Normal stress of B\u2013B A Shear stress of A\u2013A B Shear stress of B\u2013B Fig. 5 Fracture failure in section A\u2013A in Fig.\u00a02 (Figure source: literature 3) Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:273 Page 5 of 15 273 4. Torque Based on the superposition principle, maximum points of normal stress of A\u2013A and B\u2013B sections are located at point K (Fig.\u00a01). The compound normal stresses are Several observations can be made from force analysis equations: \u2022 Mp provides tensile stress for upper part of these two sections (A\u2013A, B\u2013B) and shear stress for root of planetary gear system shell", + " The model includes 178,123 elements (Solid 185 in ANSYS) and 312,564 nodes. The Young\u2019s elastic modules, yield limit, Poisson\u2019s ratio and density of housing were 200 GPa, 300\u00a0MPa, 0.3 and 7800\u00a0kg/m3. Full constraints were set on the holes of upper and lower hinged ears. A center point was applied for applying loads on the gearbox. External excitation consists of cutting resistance ( Pz ), propulsive resistance ( Px ), axial force ( A 1 ), torque ( Mp ) and gravity (G-including transmission system), as shown in Fig.\u00a02. 1. Pz and Px where NH , , n and Dc are motor rated power, mechanical efficiency of cutting part (0.85), rated speed of drum (rad/min) and diameter of drum (mm), respectively. Propulsive resistance Px is opposite to the direction of traction speed, and its value can be regarded as a linear relationship with the cutting resistance. (17)Pz= 1.91 \u00d7 107NH Dc (18)Px=kqPz Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:273 1 3 273 Page 6 of 15 where kq is a coefficient related to the wear of the cutting teeth, and its value can be regarded as 0", + " A 1 Axial force A 1 consists of propulsive resistance caused by oblique cutting and reaction force caused by coal loading. The distance between resultant point of axial forces and center of planetary gear system shell is: where LD , 0 , J are center distance between trapping shoes (mm), maximum angle between body of coal cutting machine and the straight line in oblique cutting, efficient cutting depth (mm), respectively; LC , R , K 2 are distance between trapping shoe and axis of drum (mm), radius of drum (mm) and working condition coefficient (2), respectively, as shown in Fig.\u00a02. 3. Mp According to our previous research findings [24], the amplitude of meshing force between ring gear and planetary gear is 2 \u00d7 104. (19)A 1 = DcLD sin 0 4JLc PxK2 (20)L 0 = R 4 Parameters of coal mining machine were taken to the formulas, and the results are shown in Table\u00a02. Apply these forces on finite element models, and the stress concentration regions can be obtained. Figure\u00a07 shows the von Mises stress response of stress concentration region with external excitation caused by coal cutting", + " It can be obviously seen that the vibration acceleration increases with the growth of cutting depth and traction speed. In Fig.\u00a010a, cutting depth is 300\u00a0mm before 180\u00a0 s and 600\u00a0 mm after 180\u00a0 s. Acceleration response in working condition with 600-mm cutting depth is obvious larger than 300\u00a0mm. In Fig.\u00a010b, acceleration is larger with cutting depth increases from 0 to 600\u00a0mm. Therefore, the measured vibration signals are reliable for further investigation. Frequency response analysis was conducted using the vertical accelerations of the gearbox housing in section A\u2013A (Fig.\u00a02) to investigate the local resonance characteristics of gearbox housing. Figure\u00a011 shows the power spectral 1 3 densities of vibration accelerations. It can be observed that the dominant frequencies are 950\u00a0Hz, 1250\u00a0Hz and 1400\u00a0Hz. In oblique cutting, the two dominant frequencies (1250\u00a0Hz and 1400\u00a0Hz) are close spacing, and it may cause beat phenomenon. The dominant frequencies of straight line cutting are 950\u00a0Hz and 1400\u00a0Hz, 1250\u00a0Hz is not found in straight line cutting process. However, the dominant frequencies are not necessarily the natural frequencies of gearbox housing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001543_ecce.2019.8913045-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001543_ecce.2019.8913045-Figure2-1.png", + "caption": "Fig. 2. MABE Washing Machine Motor: Topology and Stack", + "texts": [ + " The losses in the motor are fed into a one-dimensional (1D) model [2]. The model used is a 15-node thermal circuit model. It represents the washing machine motor axial and radial thermal paths [3]. The cooling features are also included in this model. The model is used to evaluate the thermal behavior of the motor during the full operation of the machine over the washing cycle. III. ELECTROMAGNETIC CHARACTERIZATION OF WASHING MACHINE MOTOR The electric motor considered for this work powers a washing machine manufactured by MABE (Fig. 2). It is a capacitor-run single phase induction motor (SPIM) operating at 50 Hz and 60 Hz. The stator has 24 slots and spans a main and auxiliary winding. The rotor has 33 aluminum bars with 1 stator slot pitch skew and aluminum casted end rings with cooling fins. The motor is rated 0.33 hp at 120 V, 60 Hz and 1690 rpm for clockwise (CW) operation. In the counterclockwise (CCW) operation mode, the motor output power is slightly less for the same input power conditions. The change in rotor rotation direction can be achieved by switching the capacitor to the auxiliary windings and reversing the main windings voltage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003307_aim43001.2020.9158857-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003307_aim43001.2020.9158857-Figure2-1.png", + "caption": "Fig. 2 Dynamic model of the series elastic actuator", + "texts": [ + " Unlike previous SEAs that used commercially available coils springs [10, 15], the costumed-made side spring has a planar structure that can be easily miniaturized to be adapted for the compact space between the screw nut and the output bracket. The side springs are made of annealed plastic mold steel (S-STAR) and the two side springs together provide a stiffness of 110.38 N/mm in both tension and compression directions. The deformation of the side springs is measured using a linear optical encoder that has a resolution of 1 \u03bcm. A rotary encoder with a resolution of 1.8\u00b0 is used to measure the motor rotation. The dynamic model of the SEA is shown in Fig. 2. For the proposed stepper motor, the rotor rotates with angle \u03b8 to translate the nut with displacement D. The linear and rotary motions are related by an amplification factor \u03b1 such that D and F (1) where F is the equivalent linear motor force. Symbol D also denotes the displacement of the reflected rotor mass Mm. The reflected rotor mass includes the moment of inertia from the motor rotor and ball screw. The mass and displacement of the output bracket are denoted as Ma and Da, respectively. The output mass Ma can further include any mass that is connected to the output bracket", + " The output and nut displacements are related by ax D D (2) where positive x denotes the compression of the side springs. The motor force F and output force Fa are both described along the screw axis. The modified motor force \u03b7F accounts for the equivalent Coulomb friction coefficient \u03b7 while the equivalent viscous friction is accounted for using a damper with a coefficient of \u03be. Both coefficients \u03b7 and \u03be are used to model the combined sliding and rolling contact of the ball screw and linear guide. Considering the model in Fig. 2, the dynamic equations governing the SEA motion can be derived [9] as mF D kx M D (3a) a a aF kx M D (3b) In Eq. (3), the motor force F is obtained using Eq. (1). The term kx is equal to the spring force f where k is the spring stiffness. The spring force is obtained given the value of x measured using the linear encoder. The value of Da is calculated using Eq. (2). The coefficients \u03b7 and \u03be have been identified experimentally. Table 1 lists the specifications of the SEA. To provide a basis for the impedance and admittance controller design, Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001588_sielmen.2019.8905795-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001588_sielmen.2019.8905795-Figure1-1.png", + "caption": "Fig. 1. Asynchronous motor used in railway traction.", + "texts": [], + "surrounding_texts": [ + "Keywords: railway asynchronous traction motors, modelling, simulation.\nI. INTRODUCTION The use of electric locomotives in railway traction is due to the high efficiency of the electric motors, the possibilities to increase their unit power, the possibilities to regulate the speed and power of the traction motors, and also to the significantly better environment conditions.\nAs the efficiency of the electric traction motors may be even greater than 93%, the electric traction is favored by the regenerative braking. In this manner, a transformation of the kinetic energy of the convoy in electricity supplied to the power line is allowed. This transformation takes place during the braking, stopping or downhill.\nWhen the regenerative braking is not possible, the motor can automatically switch to rheostatic braking. Thus, the active energy of the motor operating as a generator is transformed in heat dissipated on the braking resistors.\nThe railway traction system uses an asynchronous motor supplied by a static converter. The modern electric traction eliminates some of the existing drawbacks, by replacing the autonomous current inverters with autonomous PWM-controlled voltage inverters.\nThus, the voltages and currents provided by these inverters have waveforms very close to the sinusoid, the higher order harmonics being eliminated at the same time.\nAll of these became possible due to the occurrence of the fully controlled thyristors, high power GTOs, able to switch currents ranging from 2500 \u00f7 3000 A, supporting inverse voltages up to 4500 V.\nThese modern components significantly simplified the diagram for the load side of the voltage source inverter, eliminating the devices necessary for a forced blocking of the usual thyristors.\nThree-phase wound rotor asynchronous motors were introduced at the same time, as they are simpler and more reliable, have low overall dimensions, weights and operating costs.\nIn the modern electric traction [1-4] are used electronic converters with IGBTs, converters for ancillary services, control electronic equipment based on microcontrollers and microprocessors, LED lights, as in Fig. 2.\nIn Fig. 2 may be noticed: 1 \u2013 the electric locomotive, 2- buffers, 3 - driver's cab with all standard features, 4 \u2013 traction transformer, 5 \u2013 pantograph, 6 \u2013 brake system.\nThe interest granted to the asynchronous motor transport vehicles came along with the development of the research", + "The research is focused on finding new materials, the development of new topologies of electric motors, and also on the use of control and power electronics for increasing the effectiveness of the energy conversion.\nThe idea of obtaining the required performance from an asynchronous motor used in railway traction, such as maximum torque, minimum manufacturing and exploitation costs, is an intense concern worldwide [5-8]. These are requirements in establishing a mathematical model as precise as it can get, used in the design process, and also for the computation of the motor parameters.\nDue to the evolution of computer science and specialized software programs for solving the electrical engineering problems, the concerns related to the development of new methods and procedures for optimizing the construction of asynchronous motors used in traction systems knows a permanent evolution [9-10].\nIn such circumstances, the asynchronous motor must bear both mechanical and electrical stresses, and for a proper functioning it is necessary to know the characteristic quantities of the operating mode.\nThe multiple and repeated simulations highlight that the design is correct when, for the tests related to the particular modes, eg. idle running or short-circuit, the results meet the requirements of the beneficiary. With the measured values, one may determine the parameters for the motor\u2019s equivalent diagram.\nThe quality of the obtained simulations [10-12] depends on the used mathematical model, which must account for the basic electromagnetic phenomena that occur inside the motor. The correctness of the conducted simulations is important in operation, as the obtained parameters and results may help in anticipating the motor behavior under steady state or dynamic conditions.\nII. RESULTS OBTAINED THROUGH SIMULATION Theoretical research and experimental tests were carried out for the traction motor from a diesel electric locomotive with a power of 2100 hp. This is an asynchronous three-phase traction motor with wound rotor, with the following parameters:\n- rated values - PN=260 kW - rated power; UN=1500 V \u2013 rated voltage; I1N=121.4 A \u2013rated current; n1=1000 rpm \u2013 synchronous speed; MN=2537 Nm \u2013 rated torque; sN=2.1% - rated slip;\n- electric parameters Rs=0.247 \u03a9, R\u2019r=0.17 \u03a9, Ls\u03c3=0.00255 H, L\u2019r\u03c3=0.001818 H, Lsh=0.081 H, J=2.791 kgm2.\n- operating parameters: cos =0.90; =0.916; Mmax=2.48\u22c5M;\n- imposed gauge dimensions: Lmax < 580 - mm maximum length; Dmax < 795 mm \u2013 maximum outer diameter.\nA. Simulation of idle running performance In idle running, the slip is very small (negligible), so in the equivalent diagram corresponding to the load operation the load impedance becomes infinite. Therefore, in idle running, the motor receives from the power supply the current I=I10 and the following idle running power:\nvmFe10Cu10 pppP +++= (1)\nThe losses from the stator winding in idle running are:\n2 10110 3 IRpCu = (2)\nwhere pFe \u2013 iron losses; pm+v \u2013 mechanical losses due to the friction from the bearings and due to the ventilation.\nThe motor provides a small electromagnetic torque, needed to compensate the losses torque. The mathematical model used for analyzing the operation in idle running [13-14] is briefly presented and multiple data, obtained during the motor design process, are used.\nAt the idle running test, the motor is supplied with a voltage varying between the limits:\nfUU 110 )2,10( \u22c5\u00f7=\nFor the simulations is used the phase voltage:\n3 001.0001.0 1 110 N f UiUiU i \u22c5\u22c5=\u22c5\u22c5=\n- the induced back electromotive force is determined, with the value of the current from the previous step, I10.i-1;\n1i101i10i1e IXUU \u2212\u2212=\n- the useful fascicular flux is:\n1B11\ni1e i0h kNf2\nU\n\u03c0 =\u03a6\n- the magnetic flux density from the air gap is computed (as the gauge dimensions are in mm):\nii\nh\nl B i i \u22c5\u03c4\u22c5\u03b1 \u03a6\u22c5 = 0 6 0 10\n- the magnetic induction from the stator yoke is determined with:\n1jFeFe\ni0h 6\ni1j hlk2 10 B \u03a6\u22c5 =" + ] + }, + { + "image_filename": "designv11_80_0001299_aim.2019.8868869-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001299_aim.2019.8868869-Figure9-1.png", + "caption": "Figure 9. Assembly diagram of the amplification mechanism", + "texts": [ + " A continuous quasi-static sinusoidal waveform signal was applied to the piezoelectric actuator. The measurement results are shown in Table 1. 60 14.71 60 19.35 70 17.73 50 16.23 80 20.73 40 13.43 90 23.48 30 10.21 100 26.05 20 7.44 110 29.21 10 3.88 The hysteresis characteristic curve of the piezoelectric actuator is plotted according to Table1 and as shown in Fig. 8, and the hysteresis loop is calculated to be 14.6%. The piezoelectric actuator is coupled with the FHDA, and the structural assembly diagram is shown in Fig. 9. Whether the preload force is reasonable or not will directly affect the performance of the hinge amplifying mechanism. If the preload force is too small, the displacement of the mechanism will be lost and the rigidity of the mechanism will be reduced. If the preload force is too large, the maximum output displacement of the piezoelectric actuator will be reduced and the performance of the piezoelectric actuator can be lowered. Reasonable preload force is mainly determined by experiments. Fig. 10 shows the relationship between the preload force and the hinge output displacement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure82.8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure82.8-1.png", + "caption": "Fig. 82.8 CA GCI and AlSiC profile of von Mises stress (Model 2)", + "texts": [], + "surrounding_texts": [ + "\u2022 Suggested models of brake discs are of a solid design which needs to be pressed on to the axle before pressing of the wheels. For replacement of these brake discs, wheels are necessary to be pressed out, which is not required if the models are of a split type. Hence, further experiments may be done with split-type brake disc models. \u2022 Carbon matrix composites may be used to reduce the sound barrier and higher resistance to temperatures. \u2022 Brake pad analysis with various materials suitable to the disc material may be carried out to optimize pad material. 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 981 Appendix Structural load calculations: In this case, a railway vehicle travelling at a speed of 160 kmph on a horizontal track stops due to application of emergency brake was considered. Time of travel before stopping, deceleration, weight of the vehicle, clamping force on the brake disc, brake pad area, coefficient of friction, etc., for calculating the loads are taken from the railway specification. Value of Mass of railway vehicle\u2014M = 64000 kg, No. of axles per vehicle = 4, Maximum load per axle = 16000 kg, no. of brake discs per axle = 2, Load on each wheel = 8000 kg, Start speed v0 = 44.4 m/s, Deceleration a = 1.2 m/s2, Braking time ta = 36 s, Effective radius of the brake disc rdisc = 0.247 m, Radius of the wheel rwheel = 0.458 m, Mean coefficient of friction brake pad \u00b5 = 0.35, Clamping force Fc = 42.1 kN, Surface area of brake pads Ac = 400 cm2, Maximum temperature under sun = 70 \u00b0C, Maximum temperature under shade = 45 \u00b0C, Factor of Safety = 1.5 (Fig. 82.10). Stopping distance S \u00bc vots 1 2 at2s \u00bc 822:24m \u00f082:1\u00de Determination of pressure on disc Pressure acting on the brake disc, P \u00bc Fc Ac l \u00bc 42:1 1000 800 \u00f010\u00de 4 0:35 \u00bc 1:5Mpa on each side \u00f082:2\u00de 982 E. Madhusudhan Raju et al. where Fc Clamping force (i.e. 42 kN) Ac Contact area of brake pad on each side (i.e. 400 cm2) l Coefficient of friction (i.e. 0.35). Angular velocity x \u00bc Velocity radius \u00bc v0 rwheel \u00bc 44:44 0:458 \u00bc 97:12 rad=s \u00f082:3\u00de Thermal load The kinetic energy for one wheel (disc brake) is equivalent to the energy balance 0:125 1 2 M v2 \u00bc Zts 0 P\u00f0t\u00dedt \u00bc 2 Fdisc Zts 0 vdisc\u00f0t\u00dedt \u00f082:4\u00de The energy change at the moment is equal to the heat flux on the surface of the disc. Equation (82.4) is valid in the case of constant braking deceleration. The braking force on the disc is equal to Eq. (82.7) Fdisc \u00bc 0:125 1 2 M v20 2 rdisc rwheel v0 ts 1 2 a t2s \u00bc 8940N \u00f082:5\u00de The heat flux at the moment, which affects one half of the disc, is calculated according to the Q\u00f0t\u00de \u00bc Fdisc vdisc\u00f0t\u00de \u00bc Fdisc rdisc rwheel v0 a t\u00f0 \u00de \u00bc 8940 0:247 0:458 44:44 1:2 t\u00f0 \u00de \u00bc 214261 5786 tWatts \u00f082:6\u00de Area of friction surface \u00bc p 4 0:642 0:352 2 \u00bc 0:45m2 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 983 Q\u00f0t\u00de \u00bc 214261 5786 t 0:45 W/m2 \u00f082:7\u00de For the case of emergency braking on horizontal track from 160 kmph to stop, the analysis was carried out in 36 steps, each step being 1 s long." + ] + }, + { + "image_filename": "designv11_80_0002846_012072-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002846_012072-Figure4-1.png", + "caption": "Figure 4. Detail assembly of ball screw, load cell 2, and plates. The main plate is moved through input wheel, gear box, ball screw, secondary plate, and load cell 2.", + "texts": [], + "surrounding_texts": [ + "The based on both free body diagram and kinetic diagram shown in figure 6, the equation of friction force can be written as equation (1). (1) where Fg is friction force between the tire tread and the main plate, Fd is force measured by load cell 2, Fl is friction force of the linear motion, mp is mass of the main plate, and ap is the acceleration of the main plate. If friction force of the linear motion is negligible and the velocity of the plate is constant, the force measured by load cell is equal to the friction force between the tire tread and the main plate. ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10.1088/1757-899X/788/1/012072" + ] + }, + { + "image_filename": "designv11_80_0001748_s1052618819060074-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001748_s1052618819060074-Figure1-1.png", + "caption": "Fig. 1. (a) ABB IRB Flex-Picker robot based on Delta structural scheme, (b) amateurish Tevo Little Monster 3D printer based on the Delta scheme with linear drives.", + "texts": [ + "3103/S1052618819060074 Parallel mechanisms are used in various fields of technology: carrying and lifting equipment, machines, medical robotic instruments, and many others [1\u20133]. One of the most widespread and universally used mechanisms is the Delta robot developed by R. Clavel [4]. Developed specifically for packingand-sorting lines of food production, the Delta-robot is designed for high-speed technological operations for manipulating products with a relatively low weight (often up to one kilogram). An example of a device based on the structural Delta-scheme is the industrial robot of ABB Company of the IRB Flex-Picker series (Fig. 1a). In the classical version, the mechanism has three kinematic chains with rotational drives and intermediate parallelograms that ensure three translational degrees of freedom to the output link. In addition, a telescopic central link is used, which makes it possible to rotate the operating element. In the last few years, a variant of the mechanism with three linear drives (Fig. 1b) gained enormous popularity due to widespread use in 3D-printers. The output link has three translational degrees of freedom. Using simple gears for transformation of the rotational motion of relatively cheap stepper motors into a translation one (screw\u2013nut, belt transmission), together with the possibility of using an extrusion aluminum profile as guides, has made it possible to reduced significantly the price of the final product and helped printers implemented by this scheme to gain a large share of the market of 3D printing devices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure5.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure5.10-1.png", + "caption": "Figure 5.10 Scheme of an underground tunnel structure.", + "texts": [ + " In that case, the normal stresses are greater than in the case of a deformable ring. With an increase in the flexibility of the ring, these stresses decrease and in the extreme case, of an absolutely flexible ring they approach zero. 5.5 Calculations of Underground Structures with Arbitrary Cross-section under Seismic Action Impact Let us examine an underground structure with arbitrary cross-section, which is located in an infinite (limitless, boundless) space under a plane strain conditions (Figure 5.10). The seismic action is represented in the form of a plane (flat) longitudinal or transverse wave. The calculations of tunnels with arbitrary cross-sections to seismic resistance are done using the method of the point satisfaction of boundary conditions. This method previously was applied at torsion of bars at a bent stability and oscillations of thin plates of complex configurations.1 As seen below we propose to use this method to solve problems of diffraction of elastic waves around obstacles of complex form" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001999_14484846.2020.1714352-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001999_14484846.2020.1714352-Figure11-1.png", + "caption": "Figure 11. The given, calculated and SimMechanics conical helix trajectories of the tool tip of the 4RSS+PS PM.", + "texts": [ + " h\u00f00\u00de \u00bc 55 cm; \u03c6x\u00f00\u00de \u00bc \u03c0 6 rad; \u03c6y\u00f00\u00de \u00bc 0 rad; \u03c6z\u00f00\u00de \u00bc 0 rad Initial conditions of actuators are calculated using the inverse position kinematic and velocity analyses (Equations (12) and (26)), as follows \u03b81\u00f00\u00de \u00bc 0:33 rad; \u03b82\u00f00\u00de \u00bc 0:63 rad; \u03b83\u00f00\u00de \u00bc 0:01 rad; \u03b84\u00f00\u00de \u00bc 1:82 rad _\u03b81\u00f00\u00de \u00bc 1:04 rad=s; _\u03b82\u00f00\u00de \u00bc 1:41 rad=s; _\u03b83\u00f00\u00de \u00bc 0:54 rad=s; _\u03b84\u00f00\u00de \u00bc 0:19 rad=s Then trajectory of the tool tip is obtained as a function of time using the developed program in Matlab software, as shown by solid blue line in Figure 11. One can easily see that the calculated trajectory have no variation from the given conical helix trajectory in Equation (80). The trajectory is also obtained by SimMechanics model of the forward dynamics (Figure 11) which is so close to the given trajectory. This verifies correctness of the presented dynamic modelling of the 4RSS+PS PM. This paper dealt with the inverse and forward dynamic analyses of a novel four DOFs 4RSS+PS PM. The moving platform of the PM has a 3R1T motion. Unlimited rotation of the moving platform around the axis of PS leg and its translational motion along the same axis make the manipulator so suitable for machining applications in holes, cylinders and any places requiring an axial rotational motion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003291_012088-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003291_012088-Figure1-1.png", + "caption": "Figure 1. The vibration signal acquisition experiment.", + "texts": [ + "1088/1757-899X/892/1/012088 2 1 ( ( ) ( )) R ij i j i r j r r D E E m m (7) Then, the supporting degree ( )iS E of all evidences to iE is defined and normalized: 1 1 1 ( ) ( ) ( ) ( ) 1 \uff0c N N i in i i n n n S E D S E S E S E N (8) Then, the belief weight i of evidence iE can be obtained, and the mapping relationship between ( )iS E and i is as follows [10, 16]: ( )(1 ( ))e iS E i iS E (9) The belief weight of all evidences can be obtained based on above formula, and the basic belief functions can be modified, and the modified basic belief function of evidence iE can be expressed as: 1 ( ) ( ), ( ) 1 ( ) i r i i r r N i n n r r n m m m m (10) In this paper, the vibration signals of different fault degrees of gears are collected in DDS fault simulation test bench, and the proposed method is verified based on this. The vibration signal acquisition experiment is shown in Figure 1, and the simulated gear fault degree is shown in Figure 2. Several vibration sensors are used to collect the vibration signals of different gear fault degree, and the motor speed in set to 45 Hz in this experiment. During the experiment, for each vibration sensor, 80 samples were obtained for each gear fault degree, of which 60 samples are training samples, and there are 300 training samples in total. The remaining 20 samples are testing samples, and 100 testing samples in total. In this paper, the vibration signals of 3 vibration sensors are used, but due to space limitations, the vibration sensor 2 on the top of gearbox is taken as an example in the following" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000546_978-3-030-20131-9_278-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000546_978-3-030-20131-9_278-Figure1-1.png", + "caption": "Fig. 1. The spinal column (a) and initial movements: (b) flexion/extension, (b) left/righ lateral flexion and (c) left/right axial rotation.", + "texts": [ + " The paper presents a novel 3 DOFs waist mechanism that is used in redesign of the existing solution of multi-segment lumbar structure [2] of the social humanoid robot SARA. Some of the features of this robot are shown in [3-6]. The redesign goal is to extend the assortment of the movements and the motion range of the robots trunk. * The waist mechanism described in this paper is patent pending. T. Uhl (ed.), Advances in Mechanism and Machine Science, Mechanisms and Machine Science 73, https://doi.org/10.1007/978-3-030-20131-9_278 2809\u00a9 Springer Nature Switzerland AG 2019 Figure 1a shows the multi-joint structure \u2013 spinal column, formed from a large number of short and different bones of irregular shape \u2013 vertebrae, between which there are viscoelastic elements \u2013 intervertebral discs. The joint structures of the spinal column are the cervical, thoracic and lumbar region. Mobility between the two adjacent vertebrae is very small \u2013 only a few degrees, however the mobility of the spinal column as structure is large, as it represents the sum of a large number of short movements generated in 23 movable joints [7]. The basic movements of the spinal column are flexion/extension or bending of the body forward/backward \u2013 Fig. 1b, then left/right lateral flexion or side bending of the body \u2013 Fig. 1c, and left/right axial rotation of the body \u2013 Fig. 1d. The largest mobility is in the area of the cervical spine, while it is smallest in the thoracic region. According to [8], the range of motion of the lumbar spine of an average adult male is 40\u00f773\u00b0 for flexion movements and 7\u00f729\u00b0 for extension movements. Left and right lateral flexion movements are approximately equal and amounts 16\u00f728\u00b0 and 15\u00f728\u00b0, respectively, while the range of left and right rotation movements is identical and amounts \u00b17\u00b0. Based on [9], the maximum velocity of the lumbar region movement for males aged 20 to 29 years, at body ascent \u2013 returning of the body to its initial position, are 51", + " The rotation angle of link 8 relative to link 6 is determined according to: ( ) ( )22 2 0 9 2 0 9 16 8,1/ 2 1 2 2 arctan M L s LO M L s A A LO \u03b8 \u2212 \u2212 \u00b1 + \u2212 \u2212 = + (15) where: ( )22 2 2 2 0 9 2 1 22 O L LM s O N NM A O N + \u2212 + \u2212 = (16) The angular velocity of link 8 relative to link 6 is determined according to: ( ) 6 9 8 6 6 2 8 8 10cos sin cot s O N \u03b8 \u03b8 \u03b8 \u03d5 = \u2212 (17) where: 6 2 8 2 10 cos arccos O N O L NM \u03b8\u03d5 \u2212 = (18) Based on the kinematic analysis of the waist mechanism, the kinematic model was formed and motion simulation for initial movements of flexion, extension, lateral flexion and rotation \u2013 see Fig. 1b, using MATLAB software is performed. Input links motion ranges of the waist mechanism are determined by inverse kinematics, however the procedure is not shown because it exceeded the scope of this paper. The adopted motion law of driving links is a fifth-order polynomial, and the duration of all movements is 1s. Considering the dimensions of the robots lumbar spine, in Table 1 are shown the input geometric parameters of the waist mechanism. Figure 4 shows motion simulation results of the waist mechanism as time histories of the output kinematic parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003239_tmag.2020.3012200-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003239_tmag.2020.3012200-Figure9-1.png", + "caption": "Fig. 9. Torque in the counterclockwise.", + "texts": [ + " The force constant of d-phase at the center is zero, but it is increased according to the moving distance of the mover. In addition, the torque constant of dphase is the largest at the center. Fig. 7 illustrates the result of the thrust in the X-direction. Each phase current is determined as ia=0A, ib=-0.71A, ic=0.71A, and id=0A. Fig. 8 indicates the result of thrust in the Y-direction. Phase current to generate the thrust in Y direction is applied as ia=0.82A, ib=-0.41A, ic=-0.41A, and id=0A. Finally, the result of the torque in the counterclockwise direction is shown in Fig. 9. Currents are ia=0.29A, ib=0.29A, ic=0.29A, and id=-0.87A. From the FEA result, we confirmed that the proposed actuator can oscillate as two-DOF linear oscillatory motion and one-DOF rotational oscillatory motion. B. Dynamic Simulation The main conditions of the dynamic simulation are shown in Table I. TABLE I SIMULATION CONDITION Parameter Symbol Value Unit Mass of mover m 0.06 kg Moment of inertia J 1.2 \u00d7 10-5 kg\u00b7m2 Detent Fd Look-up table N Friction load at linear motion Fl 0.5 N Stiffness of spring ks 8000 N/m Angular restoring constant k 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001319_memsys.2019.8870648-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001319_memsys.2019.8870648-Figure7-1.png", + "caption": "Figure 7: Finished lens without final cover. The side length of the black case is 10 mm", + "texts": [ + "2 mm needles through the lower vertical segment of the lens, one for injecting the fluid and the other to release air and to remove air bubbles and excessive fluid using vacuum as shown in fig. 6. To monitor the filling process, potential air bubbles and the pre-displacement, we observe the lens with a microscope through a beam splitter, through which we can also couple a laser beam to monitor the lens curvature. To seal the lens, we finally glue it into its case, bonding the whole lower vertical segment of the elastomer structure to the case. The finished lens with an FPC connection (without the final cover) is shown in fig. 7. To characterize the performance, we measured the membrane deformation using an optical surface profilometer. This scans the surface pointwise during periodic actuation with a chromatic confocal sensor to reconstruct the dynamic displacement. As the fluid has no birefringence and the strains in the glass membrane and the polyurethane are below 10-4, this is equivalent to an optical wavefront characterization. In fig. 8, we show the focal power in the bending and buckling modes, measured in the quasi-static limit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001185_10402004.2019.1664685-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001185_10402004.2019.1664685-Figure4-1.png", + "caption": "Figure 4. Kinematical model of fluid action in a no-load multiplate wet clutch.", + "texts": [ + " distribution of the clearance flow field will be changed, which will, in turn, affect the motion of the friction pairs. Therefore, there is a coupling relationship between the friction pair and the clearance flow field. To determine the motion of the friction pair, the fluid force and torques acting on the friction pair must be analyzed. Taking the flow field formed by the FP-A and SP-a as an example, a fluid action analysis model of the friction pair in a high-speed multiplate wet clutch is established, as shown in Fig. 4. Dynamic coefficients of stiffness and damping of the flow field According to the kinematics shown in Fig. 3, the transient fluid force and torques are given as Fz \u00bc \u00d0 2p 0 \u00d0 ro ri prdrdh Mx \u00bc \u00d0 2p 0 \u00d0 ro ri pr2 sin hdrdh My \u00bc \u00d0 2p 0 \u00d0 ro ri pr2 cos hdrdh , 8>< >: [1] where p denotes the fluid pressure. The transient fluid force and torques acting on FPs/SPs are time-varying and dependent on the position of friction pairs. They are derived as the sum of steady-state and perturbed components by using a first-order approximation of Taylor expansion (Hu, et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003518_mercon50084.2020.9185241-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003518_mercon50084.2020.9185241-Figure5-1.png", + "caption": "Fig. 5. Loading and boundary conditions.", + "texts": [ + " Then the specimen was kept undisturbed for 1 hour, which was sufficient to obtain a quasistationary state for the fold, on a smooth surface until it reached its stress-free state by self-opening to its neutral position as in Fig. 4(f). Same procedure was followed for specimens with thickness of 25 \u00b5m, 50 \u00b5m and 75 \u00b5m. Two edges of each specimen was attached to a rigid plate to keep one quadrant of the unit cell in plane allowing the membrane to open freely. Specimen was supported 3 mm away from the fold-lines along both edges as shown in Fig. 5, in order to eliminate the disturbances due to boundary conditions. Low stiffness adhesive tape was used to attach the specimen to the steel plate in order to minimize the moment transfer from the support. Specimen was pulled in consecutive loading 31 Authorized licensed use limited to: Middlesex University. Downloaded on October 20,2020 at 12:53:48 UTC from IEEE Xplore. Restrictions apply. steps by loading small beads to a light-weight cone attached to a string running through two pulleys as shown in Fig. 6. A nylon string was connected to the bottom edge of the membrane aligning with the crease intersecting point as shown in Fig. 5. A light-weight cone which has a weight of 0.0032 N was attached to the other end of the string for loading beads weighing 0.0064 N and 0.0010 N. The string was passed through two pulleys before connected to the cone in order to minimize the effect of swinging and vibration of the thread while loading. Images of the specimen at each loading step were taken via a digital camera directly focused on the specimen. Coordinates of the edge of the membrane at the point where the string was attached were extracted by processing the images using an in house Matlab script" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001699_1350650119893896-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001699_1350650119893896-Figure1-1.png", + "caption": "Figure 1. (a) Structure of hole-entry hybrid journal bearing. (b) The inner bearing surface.", + "texts": [ + " In the next section, the fluid film force is obtained by solving the Reynolds equation with the finite element method (FEM) and the stability of the shaft is determined by the shaft center orbits under various initial points. Then, the following section demonstrates the differences between the linear and nonlinear orbits at an identical working condition. The nonlinear stability boundaries are investigated in the following section. Finally, conclusions are given. The structure of the hole-entry hybrid journal bearing and the inner bearing surface are shown in Figure 1(a) and (b), respectively. The geometry is a typical hybrid journal bearing which has two rows of orifices and 12 holes in each row. The bearing aspect ratio is set as l \u00bc D=L \u00bc 1, and the location of the holes is defined as a=L \u00bc 0:25. Figure 2 is the coordinate system of the hybrid journal bearing system. Non-dimensional film thickness can be derived as19 h \u00bc 1 xJ cos zJ sin \u00f01\u00de where xJ \u00bc xJ=h0, zJ \u00bc zJ=h0, xJ, zJ\u00f0 \u00de are the shaft center coordinates. The dynamic viscosity of the lubricant is largely a function of the temperature. As a cooling system is often installed in the rotor-bearing system to take away the heat caused by the rotating shaft, the temperature can be considered unchanged. Correspondingly, the variation of the dynamic viscosity can be neglected. The non-dimensional Reynolds equation governing the film pressure of Newtonian lubricants is given by20,21 @ @ h3 12 @ P @ ! \u00fe @ @ h3 12 @ P @ ! \u00bc 2 @ h @ \u00fe @ h @ t \u00f02\u00de where and are the non-dimensional axes of the oil film as shown in Figure 1(b), P is the pressure, is the speed parameter, and t is the time. To obtain the pressure distribution of the oil film, equation (2) is solved by the numerical technique. In the present study, FEM is applied. The lubricant flow field is divided into four-node isoparametric elements as shown in Figure 3. The lubricant field was divided into 35 grids in circumferential direction and 12 grids in axial direction, respectively. In each element, pressure at the point ( , ) can be approximately expressed as P , \u00f0 \u00de \u00bc X4 j\u00bc1 NjPj \u00f03\u00de The interpolation function can be chosen as22 N1 \u00bc 0:25 1 \u00f0 \u00de 1 \u00f0 \u00de N2 \u00bc 0:25 1\u00fe \u00f0 \u00de 1 \u00f0 \u00de N3 \u00bc 0:25 1\u00fe \u00f0 \u00de 1\u00fe \u00f0 \u00de N4 \u00bc 0:25 1 \u00f0 \u00de 1\u00fe \u00f0 \u00de 8>>>< >>>: \u00f04\u00de Substituting equation (3) into equation (2), the modified equation can be expressed as @ @ h3 12 @ @ X4 j\u00bc1 Nj Pj " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001669_icems.2019.8921899-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001669_icems.2019.8921899-Figure2-1.png", + "caption": "Fig. 2. Mesh generation. (a) solving domain. (b) end structures", + "texts": [ + "1, which includes stator iron core, stator end winding, stator clamping finger, stator inner clamping plate, magnetic shield, stator outer clamping plate, rotor iron core, field winding and rotor shaft. There are two clamping fingers on each stator tooth, and the stator inner clamping plate is pressed above the stator clamping finger. The magnetic shield is between stator inner and outer clamping plates, and the material of which is the same as the stator core. The stator clamping finger, stator inner clamping plate and stator outer clamping plate are all non-magnetic and conductive. The mesh in the solving domain is shown in Fig. 2. For the correct calculation, the mesh refinement is adopted in the stator winding domain and the air domain. The stator and rotor iron cores, field winding, stator end winding and the magnetic shield are considered as the non-eddy-current region. The stator clamping finger, stator inner clamping plate and stator outer clamping plate are considered as eddycurrent region. The rated current and rated field current are applied to the stator winding and field winding, respectively. The leakage magnetic field for the end region can be obtained by the three-dimensional transient time-stepping finite element method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure4-1.png", + "caption": "Figure 4. Watt mechanism.", + "texts": [], + "surrounding_texts": [ + "The most common computational tool used for design and analysis of multi-body mechanical systems is ADAMS [45, 46]. Various types of links and joints along with their properties like mass, location of center of mass, moment of inertia and degrees of freedom can be assigned for a mechanism in this tool [18]. The joints defined by this tool are ideal as they do not have any clearance or deformation. Therefore, to compare performance of the mechanisms with joint clearances, clearances at joints were made in ADAMS. The clearance at revolute joint between coupler and slider has been made by making a cylindrical hole in the slider cube and attaching a cylinder to the end of the coupler. So, we can change the size of the hole to set a clearance size. Similarly, the clearance at revolute joint between crank and coupler has been made by attaching a cylinder to the end of the crank and making a hole in the attached cylinder. 3.1 Input factors Various input parameters for modeling of mechanism in ADAMS are as follows: 3.1a Clearance size: For a standard journal-bearing of journal diameter 20 mm, the clearance size ranges from 0.02 mm to 0.08 mm. However, due to wear during operation and other environmental factors, the clearance can increase. So, in this research work, the clearance size has been taken in the range of 0.02 mm to 1 mm. 3.1b Crank speed: To cover a wide range of cases, the speeds range from 100 rpm to 3000 rpm. 3.1c Contact conditions: For the modeling of contacts, ADAMS uses the contact method based on the impact function: IMPACT-Function-Based Contact. In this method, the solver computes the contact force from the IMPACT function available in the ADAMS function library. The normal force of the contact has two components: rigidity and viscous damping. The component of rigidity is a function of the penetration d. The component of the viscous damping is a function of the speed of penetration. In this model the normal force of contact is given as: FN \u00bc Kdn \u00fe STEP d; 0; 0; dmax;Cmax\u00f0 \u00de _d; d[ 0 0; d 0 \u00f07\u00de 3.1d Value of K: The revolute joint is a case of contact between two cylinders (one inside the other). So, the contact should start with a line contact and then become a 2D rectangular contact. But the value of K in this case will not only depend on the material and geometrical property but also the stress distribution between the cylinders, which cannot be determined accurately unless it is a static case and the force is applied externally. So, the researchers solved this problem by stating that the line contact in the revolute joint will only be present for two cylinders aligned with extreme precision. Also, a uniform force distribution over the length of the joint is not possible in real life conditions. Moreover, the forcedeformation diagrams for both spherical and cylindrical impact force models were studied in the literature [1\u20134] and it was found that the spherical and cylindrical forcedeformation diagrams are reasonably close. Based on these studies, we used the Hertzian contact force law between two spheres with the different parameters defined in Eq. (8). K \u00bc 4 3p hi \u00fe hj R1=2;R \u00bc RiRj Ri \u00fe Rj ; hk \u00bc 1 v2k pEk ; k \u00bc i; j \u00f08\u00de Ri, mi and Ei represent respectively the radii of the cylinders, the Poisson\u2019s ratio and the modulus of elasticity for element i. For clearance = 0.02 mm, journal radius = 10 mm and bearing radius = 10.02 mm E = 2.07*105 N/mm2; m = 0.29 Putting these values in the equation we get K = 3.37*105 N/m1.5 Similarly, for clearance = 0.1 mm, K = 3.377*105 N/m1.5 For clearance = 0.5 mm, K = 3.4*105 N/m1.5 3.1e Value of n:The value of n is usually taken to be 1.5 for metallic contacts. So, n = 1.5. 3.1f Value of damping coefficient: In ADAMS, the instantaneous damping coefficient is a cubic step function of the penetration given as: STEP d; 0; 0; dmax;Cmax\u00f0 \u00de \u00bc 0; d 0 Cmax d dmax 2 3 2 d dmax ; 0\\d\\dmax Cmax d dmax 8 >>< >>: \u00f09\u00de The value of Cmax should be approximately 1 percent of the value of K. dmax = 0.01 mm 3.2 Output factors Two factors were used as parameters for comparing the kinematic performance of different mechanisms, either displacement of the slider or angular rotation of the rocker attached to ground." + ] + }, + { + "image_filename": "designv11_80_0003075_9783527343829.ch12-Figure12.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003075_9783527343829.ch12-Figure12.2-1.png", + "caption": "Figure 12.2 Schematic illustration of different proposed extracellular electron transfer mechanisms through an anodic biofilm. The so-called \u201cdirect electron transfer\u201d (DET) along the redox enzymes anchored to the cell membrane that is in contact with the conductive surface, \u201cmetallic-like\u201d conductivity via conductive pili, redox conduction via electron hopping between extracellular redox proteins, and mediated electron transfer based on endogenous or exogenous dissolved redox molecules. Source: Schr\u00f6der and Harnisch 2017 [26]. Reproduced with permission of Elsevier.", + "texts": [ + " All these elements have been implicated in the extracellular electron transport (EET) between microorganisms and conductive surface, but their explicit roles in the process remain unclear. Electrons not only need to be transferred outside the cell but also transported over a long-distance conductive matrix for a distant (up to tens of micrometer) terminal electron acceptor (i.e. further than the typical length of a single cell). Different mechanisms have been proposed: \u201cmetallic-like\u201d conductivity along microbial nanowires, \u201credox conduction\u201d by electron self-exchanging between immobilized redox cofactors, and soluble redox mediators (Figure 12.2). For the later part, the discussion will mainly consider Geobacter or Geobacter-dominated anodic electroactive biofilms (EABs) as the example. The mechanisms of electron transfer from electrode to the microbes are unclear, while the microbes to electrode mechanisms are more understood. Electrons generating from intracellular metabolism are transferred from the inner cell across the cell membrane and then transported along the extracellular conductive matrix (e.g. via proposed metallic-like conductivity or redox conduction) until they reach the conductive surface [27, 28]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002846_012072-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002846_012072-Figure3-1.png", + "caption": "Figure 3. The different view of the experiment set-up design.", + "texts": [], + "surrounding_texts": [ + "The development concept of the plate brake tester is shown in figure 1. A braking test is carried out statically. The vehicle wheel is placed above a moving plate. The plate is moved when the wheel is in a braking condition. A force sensor is installed on the plate. Therefore, force measured by the sensor is the braking force. A braking efficiency can be calculated by dividing the maximum braking force with normal force. With this device, it is possible to measure not only braking efficiency, but also tire characteristics. This can be done if braking force resulted from vehicle braking system is strong enough to make the wheel in a locked condition. In other words, the wheel does not rotate while the plate is moving. However, more sensors will be needed to measure parameters representing tire characteristics. The experiment set-up design to do a tire test according to the concept in figure 1 is shown in figure 2, 3, 4, and 5. It requires two load cells and two LVDT (Linear Variable Differential Transformer). Load cell 1 is used to measure normal force while load cell 2 is used to measure friction force. LVDT 1 is used to measure the contact patch and LVDT 2 is used to measure plate displacement. The main plate ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10.1088/1757-899X/788/1/012072 is moved manually from the input wheel. Power from input wheel is transmitted to the main plate through gear box, ball screw, secondary plate, and load cell 2. Data needed to model tire characteristics based on Julien\u2019s Theory are , tangential stiffness, contact patch length, coefficient of friction, normal force, and slip at maximum friction force. However, the data obtained from the experiment set up are the force measured by load cell 2 against plate displacement, normal force, and displacement measured by LVDT 1. Hence, a derivation of some equations to analyse the data is needed. Those equations are used to calculate friction force based on load cell 2 reading, contact patch length based on LVDT 1 reading, and slip based on the main plate displacement. After that, an algorithm of data analysis program is made based on equations which have derived previously and Julien\u2019s Theory. ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10.1088/1757-899X/788/1/012072" + ] + }, + { + "image_filename": "designv11_80_0000413_978-3-030-17134-6_12-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000413_978-3-030-17134-6_12-Figure10-1.png", + "caption": "Fig. 10 Product visualisation (Tobias Zerger)", + "texts": [ + " Here, the order of the courses is more flexible, which allows the students to arrange their own study-programme, incorporating elective courses, industrial design internships and international student exchanges. Alongside those elements, the suggested industrial design specialisation covers some associated engineering courses like systems engineering, mechatronics and reverse engineering.Themain focus however remains onpractical designprojects with accompanying workshops and lectures on user-centred design, product-service design, human-machine interfaces, information visualisation, product visualisation (Fig. 10), aesthetic freeform CAD modelling (Fig. 11), design research and product experience. In sum, there are four large full-semester industrial design projects and four smaller semi-semester projects in the final three years, including the graduation project and thesis. All design projects are supervised by one or more academic staff members in frequent consultations. The students work in student design studios in the faculty buildings, and have access to a student workshop and maker space. All projects are documented according to academic standards" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003990_ever48776.2020.9243015-Figure19-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003990_ever48776.2020.9243015-Figure19-1.png", + "caption": "Fig. 19. Stator structure and winding configuration of the prototype motor. (a) Stator structure. (b) Winding configuration.", + "texts": [ + " However, the difference between M2 and M3 is less than 0.5%. No matter how long the lamination axial length is, M1 always has the lowest efficiency. Therefore, with the increase of stator lamination axial length, the influence of long end-winding axial length can be reduced in terms of torque density and efficiency. VI. EXPERIMENTAL VALIDATION A 180 krpm, 450W, 6-slot/2-pole HSPM motor with 2 coil-pitch windings has been optimally designed. The stator structure and winding configuration of the prototype motor are shown in Fig. 19. For high-speed operation, the rotor-bearing system, air duct system, and house of the prototype motor are designed and shown in Fig. 20. The FEM and measured phase back EMF waveforms of the prototype motor are compared in Fig. 21, and they have a good agreement. The current and voltage waveforms at 178 krpm are shown in Fig. 22. Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 22,2020 at 14:27:45 UTC from IEEE Xplore. Restrictions apply. Three 6-slot/2-pole HSPM motors with 1, 2, and 3 coil-pitch windings have been optimally designed by FEM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002532_ropec48299.2019.9057139-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002532_ropec48299.2019.9057139-Figure1-1.png", + "caption": "Fig. 1. Structure of bearing 6205, its properties and main parts.", + "texts": [ + " The mechanical frequency fm of vibration of some damage studied can be determined by (1), while the current frequency in the side bands corresponding to the damage is obtained from (2). fo = Nb 2 fr ( 1\u2212 Rb Rc cos\u03b2 ) fi = Nb 2 fr ( 1 + Rb Rc cos\u03b2 ) fb = Rc 2Rb fr ( 1\u2212 [ Rb Rc cos\u03b2 ]2) fc = 1 2fr ( 1\u2212 Rb Rc cos\u03b2 ) fm (1) fbf = |fs \u00b1 kfm| (2) Where Nb is the number of balls in the bearing, fr is the rotation frequency, Rb is the radius of the ball, Rc is the pitch radius, \u03b2 is the contact angle between the ball and the ball bearing races (see Fig. 1), fo, fi, fb and fc are the mechanical frequency due to outer raceway fault, inner raceway fault, ball fault and cage respectively, fs is the frequency of the net, k is a natural number, fbf is the electrical frequency observed as side bands of the BF. As a main objective in this work, through current signals (MCSA), a comparison is made between methods that use spectral analysis (SA) and those that work with two types of GoFTs to evaluate their accuracy under different operating conditions. The GoFTs used are KS and Chi-Square (CHI2)", + "1 mm diameter hole and a 1.6 mm diameter hole. Damage to the ball is done by drilling a ball through the cage, with a diameter of 1.6 mm, the fourth type of damage consists of a bearing exposed to corrosion, extracted at the end of its useful life from an engine that was used in the industry; therefore it is considered a distributed damage. The damage of 3.1 mm can be considered as a quite critical damage since the radius of the ball is 4 mm. The specifications of this type of bearing according to Fig. 1 are Nb = 9, Rb = 4 mm, Rc = 19.5 mm. The IM is energized from the power line (220 V/60 Hz). The Fig. 3 shows the damage used and the Fig. 4 shows the test base used to generate the database. A total of 100 signals are obtained by storing 614.4 cycles of the current signal, distributed in 65,536 samples; the signals of interest are obtained when the motor speed has reached the quasi stationary state, so from the 100 signals 2000 signals of 2560 samples each are obtained in this state. The current signal is acquired through 3 ACS758LCB-050B current sensors, conditioning the signal through a gain and DC-offset stage using the OP177; the ADS7841P converter digitizes the 3 phase currents and sends them by SPI protocol to a DILIGENT GENESIS 2 card that uses an FPGA of the Xilinx family" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001455_012165-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001455_012165-Figure4-1.png", + "caption": "Figure 4. Displacement ( Max 0.003923 mm)", + "texts": [], + "surrounding_texts": [ + "IOP Conf. Series: Earth and Environmental Science 343 (2019) 012165 IOP Publishing doi:10.1088/1755-1315/343/1/012165" + ] + }, + { + "image_filename": "designv11_80_0002614_j.promfg.2020.04.235-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002614_j.promfg.2020.04.235-Figure3-1.png", + "caption": "Fig. 3. (a) Tool structure for manufacturing of a hybrid bevel gear, (b) structure of the semi-finished product", + "texts": [ + " To compute the coefficients A, m1, m2, m3, m4, m5, m6, m7, m8, the GRGnonlinear optimisation algorithm was used [13]. The resulting stress-strain curves for tested temperatures as well as strain rates and the resulting coefficients for the flow curve approach are discussed in chapter 4.1. In this study, a single-stage tool is used consisting of an upper die and a lower die. The semi-finished product is positioned between the upper and lower die. A sectional view of the used numerical model is depicted in Fig. 3 (a). The semifinished product is divided into three parts. The core of the semi-finished product is modelled as a cylindrical deformable body and is extended by two layers which are also modelled as a deformable body (cf. Fig. 3 (b)). This numerical investigation serves to determine the thickness of the layers with the aim of an optimal material distribution in tooth body of the final hybrid bevel gear. X45CrSi9-3 is used for the first layer and 41Cr4 for the second layer. Within this numerical investigation, the flow curves of the deposition material are implemented with the analytical flow stress approach HenselSpittel-10. For the core material C22.8, data from the simufact database were used. Due to technical restrictions of the material and the welding process of the first layer is limited to a minimal initial thickness of 1 mm and the second layer to 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001298_aim.2019.8868514-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001298_aim.2019.8868514-Figure4-1.png", + "caption": "Fig. 4. A course for testing the performance of climbing up and down the stairs", + "texts": [ + " \u2022 Metrics: 1) the test shall be conducted by 1 round-trip and in 30 minutes 2) keep the records of measured total time and calculated average of moving speed based on a predetermined reference distance of the course. 3) confirm the whole robot body completely reaches start point zone and halfway point zone of the course. \u2022 An example of practical course : Mockup staircase of Naraha Center for Remote Control Technology (Figure 1) is a typical example of the test course that satisfies above-mentioned conditions. Upper figure of Figure 4 shows an outlook of the 3D CAD data of the mockup staircase. Bottom figure of Figure 4 shows a single round-trip route for testing. In this case, the course is composed by connecting two stairs sections in a reverse attitude in with a landing (landing 2). The test performers can select the inclination of the stairs from 40, 41, 42, 43, 51, 55 deg and the road width from 0.7, 0.8, 0.9, 1.0 m. The handrails are located along the road and around the landings. The objective of this test is to evaluate the robots\u2019 performance of running with dragging the cable, quantitatively. \u2022 Requirements to be satisfied: the test shall satisfy the following conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003343_j.matpr.2020.07.163-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003343_j.matpr.2020.07.163-Figure8-1.png", + "caption": "Fig. 8. Sliding mode control & the", + "texts": [ + " Discrete sliding mode control has been investigated since 1950\u2019s. The switching frequency cannot exceed the sampling frequency because the control signal u is always held constant between the consecutive sampling intervals. Because of this limitation, it results in chattering in discrete sliding mode control. Hence, a concept known as quasi sliding mode became popular due to the chattering effect (which is a zig zag like motion). The sliding mode, discrete sliding mode, band, chattering, etc. can be best understood from the Fig. 8 [22]. The chattering phenomenon, which is the inherent property of any system is always present & cannot be avoided, but can be limited. The important point is that in quasi sliding mode the sliding mode motion cannot be confined on the sliding manifold and hence the invariance properties found in continuous time sliding mode does not hold for its discrete counter-part [49]. Utkin says, in case of discrete variable structure system, the quasi sliding mode motion is possible as control action can be activated only at sampling instants. This inherently introduces a chattering known as discretization chatter and the system motion confines to a band known as the quasi sliding mode band as shown in the Fig. 8. Gao proposed a method for discrete sliding mode control using reaching law approach [50]. This reaching law assures that the system state trajectory will hit the switching manifold & undergo a zigzag like motion, thus resulting in a quasi-sliding mode motion, about the switching manifold and finally remains within a band called as the \u2018Quasi Sliding Mode Band (QSMB)\u2019. This concept of VSS is used to design the DSMC concept in our work [23]. Please cite this article as: S. M. Kusagur, G. Arunkumar and T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000000_cac.2018.8623073-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000000_cac.2018.8623073-Figure2-1.png", + "caption": "Fig. 2. Morphing schematic of aircraft", + "texts": [ + " It is produced by NextGen, an airline of USA, and aims to develop a range of research about uav such as subsonic or supersonic flight, reconnaissance, and combat [5,9]. The basic configuration parameters of the \u201cFire-bee\u201d are given in [9]. This work was supported by the grants of Aeronautical Science Foundation of China (Grant No. 20175751028). In order to affect total aerodynamic force and aerodynamic torque of the aircraft rapidly, the wing sweep of the selected \"Fire-bee\" uav can be changed continuously in the range 0\u00b0~45\u00b0, and the half of wingspan can be changed in the range 2.5m~4.5m , as shown in Fig.2. The impact on flight status and stability from morphing mainly includes the following three aspects: (i)Changes in structural parameters such as the center of gravity and the inertia of aircraft. (ii)The aerodynamic forces and the aerodynamic torque are changed with aerodynamic parameters\u2019 great change. (iii)Bring Serious inertial forces and torque disturbances. Hence, comparing with conventional aircraft, the morphing will bring new features to the modeling and control system, which needs further researches", + "0185 +1.2585 54.8806 353.7608 +46.0297 +80.4166 \u03bb \u03be \u03bb \u03be \u03bb\u03be \u03bb \u03be \u03bb \u03be \u03bb \u03be \u03bb \u2212 \u2212 \u2212 \u2212 Then Using the high order singular value decomposition method in [10], the LPV system was transformed into an equivalent polytopic LPV system (8) which has 12 vertexes. No more details here. 4 3 ( ) ( ) ( , )( , ) 1 2 2 1 1 ( ) ( )( )i j i ji j i j X A x B u\u03c9 \u03bb \u03c9 \u03be = = = + (8) To comparing the nonlinear model of morphing aircraft with the LPV model. This paper let the aircraft transition from state A to state B in Fig. 2. The simulation results are as Fig. 4. III. DESIGN OF LPV-BASED SLIDING MODE WITH SELF-ADAPTION CONTROLLER According to the analysis in the previous chapter, it can be known that the state of the attitude, speed, and altitude of the aircraft with variable length and sweeping angle undergoes great changes in the process of movement. It takes a long time to reach a new state of balanced flight. This will cause great damage to the aircraft's flight safety and flight quality, and may even cause the aircraft to lose stability and crash" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.2-1.png", + "caption": "FIGURE 7.2", + "texts": [ + "7) When the high-pressure sealing gas flows from the high-pressure side to the lowpressure side along the radial direction, the gas expands to do work; the temperature decreases and reaches the dew point temperature; the condensation of water vapor occurs easily; and the dew point temperature is related to pressure and initial saturation [2]. Fig. 7.1 presents dew temperature curves. Here, the initial dimensionless pressure is 100, and the temperature is 300K. It is clearly illustrated that the dew temperature decreases significantly with a decrease of pressure. In addition, low humidity leads to lower dew temperature. For gas with a relative humidity of Dew point temperature curves (Po5 100 and Tout5 300K). 90%, the dew point temperature is 266K when gas pressure decreases to 20, and it is 240K for a relative humidity of 10%. Fig. 7.2 shows the distribution of vapor condensation on the seal face. It can be seen that vapor condensation mainly occurs on the sealing dam region, and the vapor condensation ratio is relatively high near the exits, but the highest point is not in the inner diameter. This is because the dew point temperature changes with a change of gas pressure, and this kind of change presents a nonlinear tendency, leading to vapor condensation being more complicated. The temperature of the sealing gas film is mainly affected by factors such as the gas film thickness, seal pressure, and speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002315_012062-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002315_012062-Figure3-1.png", + "caption": "Figure 3. Even (left) and uneven (right) gear load distribution on a planetary gear and resulting bearing forces.", + "texts": [ + " In this case, tangential gear creep is more likely to be seen at the bearing (rotor- or generator sided bearing) where the \u201cpre-loadzone\u201d tooth engagement ends due to the gear\u2019s helix angle and direction of rotation. For this, the \u201cpre-loadzone\u201d tooth engagement is defined the one that a tooth passes through before rotating into the bearing load zone. In the outlook of [5], the load distribution along the gear width was identified as a highly relevant factor to be investigated, especially for multi-row bearing concepts. Figure 3 schematically depicts an even and uneven load distribution on the two gear engagements of a planetary gear. The arrows on the outside represent the resulting axial, tangential and radial forces on the planetary gear acting from the NAWEA WindTech 2019 Journal of Physics: Conference Series 1452 (2020) 012062 IOP Publishing doi:10.1088/1742-6596/1452/1/012062 sun and ring gear. The muliple arrows on the inside represent the resulting bearing load distribution. In the right pictogram, the resulting gear meshing forces are uncentered due to misalignments in the gearbox, for example tilting of the planet gear and torsion of the sun gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002692_pedstc49159.2020.9088447-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002692_pedstc49159.2020.9088447-Figure3-1.png", + "caption": "Fig. 3. Geometrical parameters of the proposed single phase DS-EFSPM generator", + "texts": [ + " The flux linkage of each coil is zero at the unaligned positions shown in Fig.2(b) and Fig.2(d). Flux paths for PM1 and PM-2 are indicated as P1, P2, P3, and P4 in Fig.2(b). The most flux lines are pass throw the stators common teeth and the rotor segments (P1 flux path). In Fig.2(b), P5 and P6 show the PM-3 and PM-4 flux paths. the average flux linkage for all coils is zero. Flux distribution in Fig.2(d) is similar to the Fig.2(b). Therefore, the rotor rotation results in the bipolar flux linkage and alternative induced (EMF). III. ANALYSIS AND SIMULATION Fig.3 shows the main geometrical parameters of the proposed S-DS-E-FSPM generator. Also, numerical values of parameters are given in Table I. The stator teeth angles ( \u22051, \u22052= \u22053) and the rotor tooth angle ( \u22055) are optimized, based on the sensitivity analysis, in terms of minimum cogging torque, maximum flux linkage first component, and minimum flux linkage Total Harmonics Distortion (THD). The inner stator PMs thickness is considered the same as the outer stator one. The fitness function is considered as follow: = \u00d7 (1) Where, Index, LF1, CogT and FLTHD are the fitness function, the flux linkage first component amplitude (mWb), the cogging torque (N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002043_iros40897.2019.8968006-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002043_iros40897.2019.8968006-Figure3-1.png", + "caption": "Fig. 3. Left: Acetabular cup fixated in the model of the pelvic bone with the defect filled up with bone substitute. Right: Femoral component fixated in the model of the porcine femoral bone.", + "texts": [ + " The computation of the generated electric field distribution in the pelvic and femoral bone, which results from the induced oscillating field in the stimulation electrodes of the acetabular cup and the femoral component, is carried out by using the software CST EM Studio2. The acetabular cup is attached to the pelvic bone, which has been altered to show a central cavity to emulate a type 2C defect, as defined by Paprosky and Magnus [14]. The cavity is replenished with bone substitute to provide an adjoining interface between the implant and the bone (see Fig. 3). Since the prototype of the femoral component is investigated with regard to a porcine femur, in which the medullary cavity is larger than in adult human femoral bones, we focused on the investigation of the influence of uncertainty in the conductivity of bone marrow and cancellous bone for the femoral component. The electric potential \u03d5(r) in the bone tissue and at the surface of the implant is computed by solving the Laplace\u2019s equation 2CST Studio Suite 2012, http://www.cst.com/. within the computational domain \u03a9 \u2207 \u00b7 [\u03c3 (r)\u2207\u03d5 (r)] = 0, r \u2208 \u03a9 (1) which represents a volume conductor with purely resistive tissue and material properties described by their electric conductivity \u03c3(r)", + " In the case, that a compartment boundary is sufficiently far away from the stimulation electrode, the electric field distribution approaches that for the homogenous Laplace equation, resulting in nullification of the influence of uncertainty in the conductivity of the corresponding bone tissue. Therefore, the electric field strength close to the stimulation electrodes is hardly affected by the uncertainty in bone tissue conductivity, which explains the minor deviations in the overstimulation volumes. However, in the upper part of the bone deviations from this shape can be found, which result from conductivity changes across the boundaries between the bone tissue compartments (Compare Fig. 3, right). In this area, the bone marrow is comparatively small with regard to cancellous bone, and the overall shape of the bone deviates from the more cylindrical shape in its lower part, resulting in two compartment boundaries in close proximity to the stimulation electrodes, which are the reason for these deviations. Regarding the uncertainty of the electric field strength in this area, the largest deviations are located in the proximity of the bone tissue boundaries. However, the overall uncertainty in the beneficial stimulation volume is approximately 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003616_j.matpr.2020.08.536-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003616_j.matpr.2020.08.536-Figure3-1.png", + "caption": "Fig. 3. DC Generator.", + "texts": [], + "surrounding_texts": [ + "A rack and pinion mechanism along with worm and wheel gear for producing electricity. The type of shock absorber used is pneumatic shock absorbers (Figs. 2\u20135). The detailed methodology is provided in Fig. 6. DC generator Rack arrangement Worm and wheel gear Pneumatic shock absorber" + ] + }, + { + "image_filename": "designv11_80_0001184_j.mechmachtheory.2019.103606-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001184_j.mechmachtheory.2019.103606-Figure10-1.png", + "caption": "Fig. 10. The 3-PRS parallel robot.", + "texts": [ + " It should also be noted that the two spherical joints attached to the moving slider are not really free to move. The allowed movement is the rotation around the line formed by the two centers of the spherical joints. Thus, it is possible to replace the two spherical joints directly with one rotational pair. The spatial inverted triangle chain becomes to a PRS limb which everyone is very familiar with, as shown in Fig. 9 (c). The fifth combination ( C \u2207 C \u2207 C \u2207 ) would become the 3-PRS parallel manipulator as shown in Fig. 10 . Similarly, the spatial upright triangle chain can be equivalent to a PSR limb. The fourth combination ( C C C ) would become the 3-PSR parallel manipulator as shown in Fig. 11 . In the previous research, the established models of 3-PRS parallel mechanism are only suitable for the special structure (e.g., Fig. 10 (a)). For the more generic 3-PRS parallel manipulator shown in Fig. 10 (b), it is difficult to establish an analytical model because the coupling relationship between its output parameters has not been revealed. However, based on its equivalent mechanism (i.e. the fifth configuration), it is easy to deal with this problem. In addition, when it is necessary to solve calibration model of the linear Delta, 3-PRS and 3-PSR mechanisms, it can be equivalent to the mechanisms presented in this paper. Then, the unified motion model established in the previous section can be differentiated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003840_3424978.3425059-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003840_3424978.3425059-Figure8-1.png", + "caption": "Figure 8: Simulation Photo of the Underwater Vehicle.", + "texts": [ + " The data acquisition and feedback system is the sensor on each section of the vehicle. They collect the position, speed, acceleration and other information of the vehicle, and transmit it to the computer [12-18]. In the structural dynamic system, the computer receives the data information transmitted by the sensor, calculates the movement instruction of the vehicle at the next moment, and then transmits it to the steering gear in each section to make the vehicle move. The Figure 7 is the schematic diagram of the underwater vehicle. Figure 8 is the simulation photo of the underwater vehicle The vehicle is controlled by the steering gear between the sections, and the movement is completed by the twisting between the sections. By receiving the displacement and speed of each section of robotic fish, the computational force is calculated by using the fluid dynamics, and then the torque of the force that should be used by each section of robotic fish at this time, and the speed at which the twisting can be carried out are obtained. The concept of machine learning of the underwater vehicle is embodied in that the vehicle can't swim by itself, so it needs to analyze the motion response under different motion conditions, and improve its action to achieve the purpose of learning swimming" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003300_j.jmmm.2020.167289-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003300_j.jmmm.2020.167289-Figure1-1.png", + "caption": "Fig. 1. The schematic diagram of sample and arc-melting process.", + "texts": [ + "2) as-cast alloys were prepared using vacuum arc-melting under argon atmosphere (Beijing Wuke arc-melting furnace). These ingots were cut as several rectangular slabs 10(X) \u00d7 10(Y) \u00d7 2(Z) mm3 by wire electrical discharge machining (where X is the direction that parallel to the bottom surface of the arc melting furnace mold, Y is the direction that perpendicular to the bottom surface of the arc melting furnace mold, Z is the direction that perpendicular to the X and Y). The schematic diagram of sample and arc-melting process are shown in Fig. 1. The phase structures of samples were measured by X-ray diffractometer (XRD, Malvern Panalytical Empyrean X-ray diffractometer). The surface morphology and component analysis of the Fe83Ga17Rx (R = La, Pr, Sm, Y, x = 0, 0.04, 0.2) as-cast alloys were examined by scanning electron microscopy (SEM, ZEISS Sigma500) and energy dispersive spectrometer (EDS). And the orientation of the Fe83Ga17Rx (R = La, Pr, Sm, Y, x = 0, 0.04, 0.2) as-cast alloys were examined by electron backscattered diffraction (EBSD, OxFord Instruments Nanoanalysis Symmetry)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003143_jomms.2020.15.291-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003143_jomms.2020.15.291-Figure5-1.png", + "caption": "Figure 5. SMP self-folding structure: Plate thickness comprising two layers of blue and dark blue of same SMP material while yellow region as self-folding part.", + "texts": [ + " This has great application prospects in various fields, including space exploration, logistics transportation and flexible electronic devices [Tolley et al. 2014]. The application of SMPs in self-folding structures has recently attracted wide attention. In this section, we design a self-folding structure using SMP materials and simulate the deformation processes of the self-folding structure from its two-dimensional plate strip to a three-dimensional structure. The initial configuration of the self-folding structure is a two-dimensional rectangular plate strip shown in Figure 5. The plate thickness comprises two layers of the same shape memory polymer material. The bottom layer in dark blue covers the whole board while the upper SMPs layer is divided into several regions of blue portions connected intermittently by smaller yellow parts of self-folding hinge regions which can later be heated locally. Figure 6 illustrates the deformation processes transforming a two-dimensional pattern to a threedimensional configuration. First, we heat the original plate structure to a high temperature and apply a compressive deformation along the length of the plate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure36.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure36.2-1.png", + "caption": "Fig. 36.2 Different raster orientations XY, XZ, YZ and ASTM D638 Type-1 tensile test specimen", + "texts": [ + " The students, researchers and even normal household users can easily install and work to produce the desired products using the different parameter combinations and of different shapes and sizes. These printers can be a great learning experience as the users can visualize, model and print in quick succession to get the rapid prototype of the specimen or product. In the present work, raster orientation and infill density are input parameters and based on their different combinations, mechanical properties like tensile strength and yield strength are calculated. The tensile test specimens were prepared based on ASTM D638, standard dimensions are used corresponding to Type-1 specimen in Fig. 36.2, and each sample set consisted of five specimens, Fig. 36.3. The geometry of the 3D-printed specimens was modelled using SolidWorks software exported as an STL file and imported to the 3D printing software. A 50 KN universal testing machine was used for tensile testing with speed of 5 mm/min, and fractography was performed using scanning electron microscope (SEM). The material used for current experiment is PLA. 418 P. Yadav et al. Before performing fractography, tensile tested samples were cut and gold (75%) and palladium (25%) coating was done for better image analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure14-1.png", + "caption": "Figure 14 Eighth mode", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0001471_j.engfailanal.2019.104223-Figure16-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001471_j.engfailanal.2019.104223-Figure16-1.png", + "caption": "Fig. 16. Assembly shift of rear fender.", + "texts": [ + " Strain gauge pasted at failure location of rear lower centre cowl. led to confirmation that pillion load transfer is not the actual cause of failure of rear lower centre cowl. In general, mounting holes of rear fender and rear cowl are oblong to accommodate frame deviations. Effect of rear fender in tolerance stack up analysis for minimum gap between rear lower centre cowl and rear fender is performed and it is found that 70% is contributed by assembly shift. The rear fender is tilted by 1.2 degree as shown in Fig. 16 due to assembly shift, thereby reducing the gap by 2mm. As per design, gap between rear cowl assembly and rear fender= h mm (without assembly shift) as shown in Fig. 17. The actual gap has been checked in tested vehicles and it is found that gap has reduced in all vehicles. With respect to actual gap measured in vehicle, the rear assembly along with frame is modelled as shell mesh with average mesh density as 2.5mm and is simulated for dynamic loading condition. Harmonic data used in electrodynamic shaker as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001192_ilt-07-2019-0264-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001192_ilt-07-2019-0264-Figure1-1.png", + "caption": "Figure 1 Schematic design, relative motion and coordinate systems of OCGDBs", + "texts": [ + " This paper presents a detailed investigation on the dynamic characteristics of OCGDBs. The steady-state Reynolds equation and perturbedReynolds equations are solved to obtain the 5-DOF stiffness and damping coefficients considering translational and tilting perturbations. Based onRouth\u2013Hurwitz criterion, a formula for quickly calculating the critical mass is deduced. Moreover the stability of the OCGDBs is analyzed by the traditional Routh\u2013 Hurwitz criterion and the formula of criticalmass. 2. Mechanical model In Figure 1(a), a schematic of a pair of OCGDBs supporting a rotor is shown. The bearings are fixed and connected by a shaft. While the rotor rotates with angular velocity v , the gas film in the clearance produces bearing force to support the rotor. Spiral grooves on the surface of the bearings can pump ambient gas into the clearance and improve the pressure. The mechanical model of the bearings is defined by bottom radius R, width b, taper kt, groove depth hg and groove angle b g. The space between two bearings is d, and the bearing clearance is c. The Cartesian coordinate systemObios is fixed on the bearings. Ob is the center of the bearings and the s-axis is the rotation axis of the rotor. The primary assumptions are listed below: the rotor and bearings are rigid; the clearance is considered to bemuch smaller than other sizes; the gas in the clearance is ideal and isothermal; and the rotor rotates at a constant speed despite disturbance. The relative motion between rotor and bearings is shown in Figure 1(b). Considering the rotary symmetry of the rotor, the position of the rotor could be described by 5-DOF motion. The new position is represented by the coordinate system Ori\u2019o\u2019s\u2019. It could be reached after translation u = (ui, uo, us) from Obios to Ori1o1s1, tilting motion around i-axis to Ori2o2s2 and tilting motion around o-axis to Ori\u2019o\u2019s\u2019. The tilting angles are w i and w o, respectively. 3. Mathematical model 3.1 Governing equations of the gas film The pressure of the gas film is governed by the Reynolds equation (Li and Duan, 2019)", + ", 2018), and the effect of gas rarefaction is described by the Frenkel\u2013Kontorova model (Fukui and Kaneko, 1990; Huang, Opposed-conical gas-dynamic bearings Yan Li, Desheng Zhang and Fuhai Duan Industrial Lubrication and Tribology Volume 72 \u00b7 Number 3 \u00b7 2020 \u00b7 415\u2013425 2012). The dimensionless transient Reynolds equation is modified for theOCGDBs in the following form: @ @u ph 3 Q @p @u 1 R=b1 ktZ kt\u00f0 \u00de2 11 k2t @ @Z ph 3 Q @p @Z \u00bc 6mv R=b1 ktZ kt\u00f0 \u00de2 Pac2 @ ph @u 1 12mv R=b1 ktZ kt\u00f0 \u00de2 Pac2 @ ph @t (1) where u and Z are coordinates shown in Figure 1(c), p is dimensionless pressure, h is dimensionless gas film thickness and t is dimensionless time. Q is defined to describe the Poiseuille flow in the Frenkel\u2013Kontorova model, obtained by the following expressions: Q \u00bc Qp=Qcon Qcon \u00bc D=6 D \u00bc D0ph D0 \u00bc Pac m ffiffiffiffiffiffiffiffiffiffiffiffiffi 2RgT0 p Qp \u00bc D 6 \u00fe 1:0162\u00fe 0:40134 ln 1\u00fe 1:2477 D 8>>>>>>>< >>>>>>>>: (2) where Qcon is the flow rate coefficient for continuum Poiseuille flow, Qp is the flow rate coefficient for Poiseuille flow, D is inversed Knudsen number and D0 is a characteristic inverse Knudsen number with the temperature T0 = 294 K and gas constantRg = 8.314 J/(mol\u00b7K). Dimensionless film thickness can be obtained by the following equation: h \u00bc 11 u1u x\u00f0 \u00de nj =c1 hg=c j \u00bc 1; 2 (3) where u = (ui, uo, us) is the displacement of the rotor, u = (w i, w o, 0) is the tilting angle of the rotor and x = (i, o, s) is the coordinates defined in Figure 1(c) and the normal unit vector nj of bearing surface points to the rotor. Subscript j = 1 refers to the cone in the positive axis and j = 2 refers to the cone in the negative axis. Figure 1(c) shows the two coordinate systems for calculation, whichmeet the following conditions: i \u00bc R1 ktbZ ktb\u00f0 \u00decosu o \u00bc R1 ktbZ ktb\u00f0 \u00desinu s \u00bc 1\u00f0 \u00dej bZ1 d=2\u00f0 \u00de 8>< >: (4) The dynamic stiffness and damping coefficients are obtained by plugging a small displacement and velocity perturbation into the Reynolds equation (Han and Fu, 2018; Lu et al., 2011). The perturbation is expressed as Dui, Duo, Dus, Dw i, Dw o, Dvi, Dvo, Dvs, Dv i and Dv o, where vi, vo, vs, are velocities along i, o, s directions, respectively, and v i, v o are angular velocities around i, o directions, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000297_kem.799.263-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000297_kem.799.263-Figure5-1.png", + "caption": "Fig. 5 LCD guidance system assembly", + "texts": [ + " In this case, modifications include, repairing ridges, cuts, groves, tears, marks and formations of built-up edges. Second, it was designed to be equipped with a coaxial powder nozzle, allowing different powder injection options. In addition, as the device is directly fitted on the crankshaft journal fillets, the device can clad in any direction following different cladding strategies (e.g. zigzag or spiral paths). To verify and collect the data regarding the mobility parameters or the capability to clad in any direction, measurements were taken using the CAD model in Fig. 5. As can be seen in Fig. 5, the LCD is equipped with six step motors, three on each side, working in parallel for each movement. This means, two motors for the lift motion (Detail B red circle), two for the transversal movement (Detail C) between the crank webs and two for the swing motion (Detail A and B) of the nozzle head. Detailed explanations about the LCD have been provided previously. Nevertheless, certain mobility aspects were improved following mechanical analysis of the LCD. In this analysis, the prototype\u2019s movement behaviour was observed, and the potential problematic areas are indicated in Table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001284_chicc.2019.8866532-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001284_chicc.2019.8866532-Figure1-1.png", + "caption": "Fig. 1: VolantexRC 757-3 FW-UAV", + "texts": [ + " What\u2019s more, in this situation, if some residual rij alarms, the neighbor UAV j will be detected to have a fault. Therefore, the condition for UAV i detecting a fault on itself is \u2016rik\u2016RMS > Jth, \u2200k \u2208 Ni (22) where Ni is the set of neighbors of UAV i. And the condition for UAV i detecting a fault on its neighbor j is \u2016rij\u2016RMS > Jth, j \u2208 Ni (23) The fault detection and isolation procedure is given by Al- gorithm 1. The proposed method has been experimentally tested using four VolantexRC 757-3 FW-UAVs (see Fig 1). The basic parameters are given in Tables 1. The parameter matrixes of the linear model (8) are given as follows Alon = \u23a1 \u23a2\u23a2\u23a3 \u22120.0144 0.7171 \u22120.94 \u22129.8 0.0864 \u22121.6845 8.08 7.42 0.0474 \u22120.6585 \u22123.72 0 0 0 1 0 \u23a4 \u23a5\u23a5\u23a6 Algorithm 1 Fault detection and isolation in UAV i Off-line parameters design: Choose observer matrix L such that \u2212L is Hurwitz. Property Symbol Value Units Mass m 2.200 kg Wing Span b 2.01 m Wing Area S 0.4087 m2 Mean Aerodynamci Chord c 0.25 m Alat = \u23a1 \u23a2\u23a2\u23a3 \u22121.2 0.5 \u221219 9 \u22122.1 \u221213 13 0 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001761_012021-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001761_012021-Figure4-1.png", + "caption": "Figure 4. Shear stress distribution", + "texts": [ + " From observations of the fracture surface using a magnitude of 100x magnification, it was observed that the initial crack occurred at the base of gear of a teeth. Moreover, It was observed that the crack propagation direction that marked with feather marks [12]. Stress and strain analysis was then conducted on the spur gear using FEM. From the results of the stress analysis on the gears, it was found that the region with the highest shear stress was at the contact surface of the gears with the value of 13,309 MPa, as illustrated in Fig. 4. URICSE Journal of Physics: Conference Series 1351 (2019) 012021 IOP Publishing doi:10.1088/1742-6596/1351/1/012021 From the simulation results can also be plotted a shear stress curve from the tooth base to pitch as shown in Fig. 5. Where the value of shear stress is taken at the node near the teeth around the crack tip. From the curve it can be seen that from the base of the tooth to tooth pitch there is an increase shear stress up to point 3 as the maximum value of shear stress. This suggests that in addition to tooth contact point where the maximum shear stress occurs, the base or roots of the teeth also hold a fairly large shear loads" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000029_s12239-019-0013-z-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000029_s12239-019-0013-z-Figure3-1.png", + "caption": "Figure 3. Wrenches acting on the coupler of a planar crank-slider mechanism.", + "texts": [ + " Hence, the twist of the coupler is found to be pointing along the z-axis and passing through the intersection of line ab and line dc; due to the three parallel wrenches acting on the coupler along the z- axis, the coupler moves in the x-y plane and the pitch of the screw is zero. The intersecting point of the instant screw axis and the plane of motion is the instant center of velocity of the coupler with respect to the ground. In the case of a planar one degree of freedom crankslider mechanism shown in Figure 3, the coupler, bce, is connected by an R-R link and an R-P link whose R axis is along the z-axis and P axis is along the x-axis. The R-R link exerts three independent wrenches W1 to W3 as shown in Figure 3. The R-P link exerts four wrenches on the coupler: a force wrench W4 at the R joint labeled c along the z-axis, another force wrench W5 at c along the y-axis, and two couple wrenches along x- and y-axes. The two couple wrenches, however, are dependent on any two of the three parallel force wrenches W1, W2, and W4 along the z-axis, and can be eliminated. Hence, the five independent force wrenches determines the zero pitch twist of the coupler relative to the ground that is along the z-axis and passes through the intersection of W3 and W5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001288_j.matpr.2019.08.229-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001288_j.matpr.2019.08.229-Figure6-1.png", + "caption": "Fig. 6. Pressure cylinder with meshing model.", + "texts": [], + "surrounding_texts": [ + "Nomenclature\nd Diameter l Length t Thickness p Intensity of max. internal pressure 1/m Poisson\u2019s ratio rc Circumferential stress\nrt Longitudinal stress dl Change in length dd Change in diameter e1 Circumferential strain e2 Longitudinal strain\nPlease cite this article as: K. Rajendra Prasad and C. Syamsundar, Theoretical applications, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.201\nof these studies were not compared with experimental approach during their analysis.\n3. Design of pressure cylinders\nAccording to the dimensions of pressure cylinders, they are mainly of thin cylinder or thick cylinder as mentioned below.\nAccording to the ratio\u2019s\nt d 6 1 10 - Thin cylinder \u00f03:1\u00de\nt d P 1 10 - Thick cylinder \u00f03:2\u00de\nIn general thick shells are mainly used for high pressures. Thin cylinders design mainly involves the calculation of thickness (t) as\nCylindrical cylinder thickness; t \u00bc pd 2rc\n\u00f03:3\u00de\ntion of cylindrical cylinder.\n(b) Sandwich reinforced composite.\nand FE analysis of epoxy composite pressure cylinder used for aerospace 9.08.229", + "Fig. 4. (a) 2-D Model of pressure cylinder, and (b) Model developed in CREO parametric.\nFig. 5. (a) Pressure cylinder model, (b) Bonded region, (c) Fixed support, and (d) Internal pressure.\nInternal pressure distribution of cylindrical cylinder is shown in Fig. 3.\nCircumferential stresses will be\nrc \u00bc Total pressure Resisting section \u00bc pdl 2tl \u00bc pd 2t\n\u00f03:4\u00de\nThe longitudinal stresses will be,\nrt \u00bc Total pressure Resisting section \u00bc pd 4t\n\u00f03:5\u00de\nNow changes in diameter and length may be found out from the above equations,\ndd \u00bc TM 1 d \u00bc pd 2tE 1 1 2m\nd \u00f03:6\u00de\ndl \u00bc TM 2 l \u00bc pdi 2tE 1 2 1 m\n\u00f03:7\u00de\nPlease cite this article as: K. Rajendra Prasad and C. Syamsundar, Theoretical and FE analysis of epoxy composite pressure cylinder used for aerospace applications, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2019.08.229" + ] + }, + { + "image_filename": "designv11_80_0001150_tasc.2019.2903652-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001150_tasc.2019.2903652-Figure5-1.png", + "caption": "Fig. 5. Structure of the measured grain-oriented electrical steel for iron core.", + "texts": [ + " In order to measure the magnetic field and magnetic flux density of the electrical steel sheets, the samples were stacked up to make an iron rod, thus it could be homogeneously magnetized. The measuring platform is mainly composed of magnetic flux density coil (B coil) and magnetic density coil (H coil), which was wound in the center of iron rod and epoxy rod. The tested electrical steel, namely iron rod, was binding together with the epoxy rod by the insulated tape. It was fixed at the center of the SC coil and formed a closed magnetic circuit with the air. The layout of the measuring platform is shown in Fig. 4 and Fig. 5. Through the use of the measuring platform, we could measure the B-H curve of the electrical steel by changing the exciting current of superconducting coils (SC coil). However, the iron core can not even improve the flux density to 2T or above as the superconducting coils reach the critical current 60A at the deeply saturated state. Therefore, we made a whole iron yoke for the sample to form a closed magnetic circuit. In this way, we can decrease the magnetic flux leakage and take full advantage of the flux excited by superconducting coils. The structure of the measured grain-oriented electrical steel for iron core is shown in Fig. 5. In the process of alternating magnetization, the magnetic field strength H and magnetic flux density B are drawn from point to point, and the curve it made is called hysteresis loop. We measured a cluster of the hysteresis loops and the in- duced voltage of the H coil varied from 50 mV to 2760 mV. As shown in Fig. 6, we can recognize that the magnetic hysteresis is not so obvious at the beginning, but it gradually appears as the magnetic field strength becomes larger. What\u2019s more, the hysteresis loops present mirror symmetry at the origin of coordinates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003442_aero47225.2020.9172494-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003442_aero47225.2020.9172494-Figure2-1.png", + "caption": "Fig. 2 Camera configuration within the body frame of a spacecraft of the formation", + "texts": [ + "\u0307?i T]T\u2208 \u211c6 contains the absolute linear and angular accelerations of the i-th spacecraft, \ud835\udc70\ud835\udc56\u2208 \u211c6\u00d76 is the mass matrix containing the mass and moments of the inertia of the spacecraft, \ud835\udc6a\ud835\udc56\u2208 \u211c6 contains the non-linear velocity/displacement-dependent terms, F\ud835\udc50,\ud835\udc56\u2208 \u211c6 contains the force and moment exerted by the satellite actuators, and F\ud835\udc52,\ud835\udc56\u2208 \u211c6 contains the external/disturbing forces and torques applied to the chaser satellite. Each spacecraft is equipped with a system of cameras, as represented in Fig. 2, where the spacecraft body reference frame and one out of six camera frames are also illustrated. The latest has the ?\u0302?\ud835\udc56 \ud835\udc50 axis aligned along the outward direction of the optical axis of the camera and the other two axes ( \ud835\udc99\ud835\udc56 \ud835\udc50\ud835\udefe and ?\u0302?\ud835\udc56 \ud835\udc50\ud835\udefe) laying on the outer panel surface of the panel. The same logic is repeated for defining each of the five other camera frames {\ud835\udc36\ud835\udefe} of the satellite. Figure YY shows the ideal projection of a point p on the image plane when it falls within the field of view of the camera" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure5-1.png", + "caption": "Figure 5. Watt mechanism with slider (1).", + "texts": [], + "surrounding_texts": [ + "ADAMS allows export of data in the form of an Excel Sheet. The data contains many data points of the graph at equal intervals of time. Thus, we have two data sets for the two curves. Now, we need to compare these data sets in order to quantify the difference between the two curves. Some statistical measures which can be used to compare the two data sets are mean deviation (obtained by taking the mean of differences in output of mechanisms with and without clearance), RMS deviation, 90% points tolerance (value of tolerance on both the positive and negative sides of the ideal curve, within which 90% of the data points of real mechanism lie), t-test (to check if the two means are reliably different from each other), F-test and KolmogorovSmirnov test. All these six measures were computed for a set of preliminary simulations with varying levels of clearance. The mean deviation was found to be the most sensitive of these measures. Thus, for the rest of the analysis, we used mean deviation as a measure of deviation of the performance of the mechanism from the ideal mechanism." + ] + }, + { + "image_filename": "designv11_80_0000025_s1068366618060144-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000025_s1068366618060144-Figure4-1.png", + "caption": "Fig. 4. Schematic of the experimental setup to study the influence of the thermal processes on the formation of the pressure in the sliding bearing: (1) rotating cylinder, (2) fixed cylinder, (3) electric motor, (4) adjusting screws, (5) f lange to secure the electric motor, and (6) heating element; (A, B, and C) points of drainage orifices for measuring the pressure in the clearance between the cylinders and the temperature sensors, (A1, A2, and A3) drainage orifices in the setup casing, and (TA1, TA2, and TA3) temperature sensors.", + "texts": [ + " 3, graphs of the excess pressure in the slitlike clearance are shown depending on the polar angle \u03d5 at different values of the eccentricity ratio \u03b4 considering changes in the temperature along the bearing axis calculated by Eq. (9). In this case, the excess pressure assumes a positive value at nodal points. Depending on the temperature and the eccentricity ratio, there may be no low-density regions in the clearance at all. Consequently, the theoretical results predict that the dissipative processes may lead to significant changes in the pressure distribution in the clearance between the cylinders. A schematic of the experimental setup is shown in Fig. 4. The setup comprises fixed cylinder 2, which serves as the casing at the same time in which electric motor 3 is secured by f lange 5. On the motor drive shaft, rotating cylinder 1 is mounted. At points A, B, and C spaced 120\u00b0 apart along the outer cylinder circumference, three drainage orifices at each point, viz., A1, A2, A3, B1, B2, B3, C1, C2, and C3, are bored along the generatrices; the orifices are designated for measuring the pressure in the clearance between the cylinders. The orifices are bored at distances from the upper edge that equal 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001658_icems.2019.8921632-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001658_icems.2019.8921632-Figure3-1.png", + "caption": "Fig. 3. Subdomain divisions of SM-MG and ST-MG: (a) SM-MG, (b) STMG.", + "texts": [ + " Thus, a twosegment equivalent method is introduced to make the magnetic field distribution of ST-MG calculatable [17]. The goal is to make the equivalent sector PM produce the same magnetic field as the rectangular PM. Thus, the PM amount should be the same. Thus, the arc angle of section 1 \u03b31, and section 2 \u03b32 can be expressed as: After the mathematical simplification, both SM-MG and ST-MG can be divided into several ringlike subdomains, which are related to one another via different boundary conditions, as shown in Fig.3. It can be seen that there are nine subdomains for both SM-MG and ST-MG. The shaft is region I; the inner air gap and outer air gap are region IV and region VI, respectively; the modulator is region V; the PM of outer rotor is region VII, the back iron of outer rotor is region VIII and the outermost air is region IX. The only difference between SM-MG and ST-MG is the inner rotor part. For SPMG, the back iron of inner rotor is region II, and the PM of inner rotor is region III. As for ST-MG, the section 1 is region II, and the section 2 is region III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001569_ecce.2019.8912739-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001569_ecce.2019.8912739-Figure3-1.png", + "caption": "Fig. 3 Beam representation of the bridges", + "texts": [ + " Finally, remarks and perspectives are presented in the conclusion. II. ANALYTICAL METHODS The study of the stress magnitudes in the rotor, for different iron bridges thicknesses, shows that the highest stresses are located inside iron bridges: the central and the two air gap bridges [5]. From a modeling point of view, critical iron bridges can be considered as simple rectangular section beams where the stack length l and thicknesses ea and ec of the air gap and the central bridges, define the dimensions of the beams as shown in Fig.3. Such a configuration allows the use of the beam theory in order to perform an analytical modeling of the mechanical stresses [5]. Each pole of the rotor is considered as embedded with the adjacent poles at the air gap iron bridges (K1 and K2), and with the rotor body at the central iron bridge (K3) (Fig.4). The centrifugal force Fcentrifugal is the main source of stress within iron bridges. It is applied at the point G, the center of gravity of the active part of mass m of the rotor which is defined as the upper part of the rotor pole including all the magnets" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001835_s12206-019-1147-7-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001835_s12206-019-1147-7-Figure1-1.png", + "caption": "Fig. 1. Schematic representation of the cart table model.", + "texts": [ + " In this paper, the model used to represent the humanoid robot when walking is a simple cart table model. Even in its simplicity, it remains a reliable model used in the literature to study ZMP motion [11]. A major reason for this choice is the capacity of such models to reflect the dynamic behaviour of the original system while staying very easy to compute for postural stability analysis [15]. The system consists in a cart of mass m situated at the height of cZ and controlled via a control input u. A schematic representation of the system is shown in the Fig. 1. The system is represented by the following continuous state space system and the ZMP problem is then translated to the following dynamical system. Linear quadratic regulators have been widely investigated for their capacity to insure tracking while minimizing the control effort [16]. In the area of locomotion, a considerable attention has been paid to such controller [7]. This attention has been motivated by the simplicity and robustness of this type of controller. Moreover, going by the principle that human walk in a way that minimizes the energy consumed, designing a controller that could achieve a stable walking with a minimum energy cost is a crucial point in the quiet of humanoid locomotion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001185_10402004.2019.1664685-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001185_10402004.2019.1664685-Figure5-1.png", + "caption": "Figure 5. Diagram of rub-impact force analysis of the friction pair.", + "texts": [ + " Rub-impact model of the friction pair In this article, the dynamic contact impact theory is adopted to analyze the impact process of the friction pairs in a highspeed multiplate wet clutch. It is considered that the impact process is microcosmic and factors such as contact force, contact time, and contact deformation are taken into account when establishing the impact model. It is assumed that the FPs and SPs always remain rigid and an impact between the friction pair happened at the dot of the outer circumference. Aiming at the anticlinal impact, the rubimpact effect of the friction pair is presented by the normal contact force Fn and tangential friction force Fs (see Fig. 5). The axial rub-impact force Fn will affect the 3 DOF motion of the friction pair and increase its degree of nonlinear motion, and the tangential friction force Fs will act as a hindrance to the rotation of the FP and increase the drag torque in the wet clutch. Once the relative axial displacement of an outer dot reaches the initial clearance h0, it is considered that the FP impacts with the SP. Considering that the friction system of a highspeed multiplate wet clutch is multiparameter, high-dimensional, and nonlinear, the nonlinear damping model is adopted" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002336_humanoids43949.2019.9035074-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002336_humanoids43949.2019.9035074-Figure6-1.png", + "caption": "Fig. 6. Telescopic Inverted Pendulum (TIP) model.", + "texts": [ + " If the constraints in (7) are always verified during the whole simulation, this can be simplified to q\u0307r = S(q\u0307T1 + q\u0307T2) (10) where S = [ 0(n\u22126)\u00d76 I(n\u22126)\u00d7(n\u22126) ] is a matrix that selects the actuated joints only. The robot can then be controlled in position, numerically integrating the velocity vector calculated in (10). To compare the results of the proposed trajectoryfollowing controller using the optimal trajectory for the CoM, we considered the online energy shaping (ES) controller proposed in [13]. The tipping humanoid has been modeled using the TIP model with the y-axis directed downward, as shown in Fig. 6. The equations of motion are: mr\u0308 \u2212mr\u03b8\u03072 +mg cos \u03b8 = 0 (11) mr2\u03b8\u0308 \u2212mgr sin \u03b8 + 2mr\u0307\u03b8\u0307 = 0 (12) The only parameter that can be controlled during the fall is the length r of the pendulum, which regulates the CoM position, i.e. the pendulum can contract or elongate in order to dissipate energy. The total mechanical energy is given by E = EK + EP = 1 2 mr2\u03b8\u03072 + 1 2 mr\u03072 \u2212mgr cos \u03b8 (13) Taking the time derivative of (13), substituting \u03b8\u0308 from (12) and neglecting the higher order terms (i.e., r\u0308) [13], the following relation is obtained E\u0307 = 2mgr\u03b8\u0307 sin \u03b8 \u2212 (mr\u03b8\u03072 +mg cos \u03b8)r\u0307 (14) Defining A = 2mgr\u03b8\u0307 sin \u03b8 and B = mr\u03b8\u03072 + mg cos \u03b8, the desired r\u0307d is found as r\u0307d = KBE +A/B (15) where K > 0 is a control gain" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001471_j.engfailanal.2019.104223-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001471_j.engfailanal.2019.104223-Figure12-1.png", + "caption": "Fig. 12. (a) Gap between Seat & Cowl in side view and (b) Impression mark left on Rear Cowl Assembly in top view.", + "texts": [ + " Warpage resulting in deviation of top mounting is countered, thereby avoiding assembly strain. However, regardless of this, failure of the part is observed at electrodynamic shaker after running for 30% of the required duration which concludes warpage in the part is not root cause for the failure. Next, the effect of surrounding components on rear centre cowl is studied. In order to avoid see through between seat and rear cowl assembly, as per styling, gap between seat and rear cowl assembly, shown in Fig. 12(a), is reduced. To check whether pillion load is transferred to rear cowl assembly, clay is placed between the two parts. Impression mark is observed in double riding Fig. 10. Weld line near the failure location. Fig. 11. Shrinkage variants causing warpage. condition, i.e. it is evident that pillion load is getting transferred to rear cowl assembly as shown in Fig. 12(b). Strain gauges are pasted at mounting region (from where crack propagated) of rear lower centre cowl as shown in Fig. 13 to acquire data for two different conditions. Firstly, for pillion load transfer in static condition of vehicle and secondly, strain data is acquired for dynamic load in electrodynamic shaker without any pillion load. Table 3 illustrates high strains at corresponding locations depicting the severity of both the events. High strain is observed during dynamic loading in shaker without pillion load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure90.12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure90.12-1.png", + "caption": "Fig. 90.12 Displacement plot in the debond model", + "texts": [ + " a void created in-between will allow liner to consume pressure load more than its capability, which will allow it to deform excessively as it cannot effectively transfer the load to composite overwrap. Peak Von Mises stress of 293 MPa (refer Fig. 90.10) is seen at the debond locations. The obtained stress of 293 MPa is 3 MPa more than the ultimate tensile strength of AA 6061-T6 proving that liner will fail in presence of debond. Graphical representation of Von Mises stress along the axial length of liner is shown in Fig. 90.11. Figure 90.12 shows the maximum deformation (7.93 mm) in liner at debond location. This plastic deformation causes 90 Finite Element Analysis of Potential Liner Failures \u2026 1083 1084 R. Pramod et al. additional problem, as during consecutive pressurization-depressurization cycles, it may cause liner to buckle; hence, bringing down the efficiency of liner and possible failure during operation. The eigenvalue linear buckling analysis is conducted to observe the buckling of liner (plastic condition) when a compressive load of 1 MPa is assigned to obtain load at which the liner buckles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003223_012006-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003223_012006-Figure1-1.png", + "caption": "Figure 1. A free body diagram of quadcopter X [9]", + "texts": [ + " This paper is to explore the relationship between a new model of a quadcopter and automatic balancing system using a combination of the mass of battery and landing gear. It is also to create a counterbalance to have a faster response and to present the process of improvements in PID control of stabilization. This section discusses the model and design of the quadcopter and concept of balancing system. The S-500 PCB quad-rotor frame was used. The quadcopter consists of four motors, there are two motors rotate clockwise (CW) direction and two motors rotate counterclockwise (CCW) direction as shown in Figure1. The moment force around the vertical axis of the aircraft body was compensated for each other. In this study, the balance situation with the equally PWM signal value of four motors used as a condition. The unbalance situation simulated by installing a payload on the front of the aircraft frame. The PWM signal value of four motors was unable to have equal values. From this data, the PWM signal values of a pair of front motors showed greater value than a pair of rear motors. The lightweight aluminum bar is attached to the front of the aircraft frame so that it was able to adjust the position of aluminum block to simulate the movement of the center of gravity when carrying a payload (see Figure 2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000816_iemdc.2019.8785225-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000816_iemdc.2019.8785225-Figure6-1.png", + "caption": "Fig. 6. Experimental setup with vibration sensors mounted close to the bearings on both the load side and the opposite side.", + "texts": [ + " Validation of the Proposed Signal Processing Technique To demonstrate the effectiveness of the proposed signal processing technique, an accelerated simulation is performed on bearing degradation, and the resultant mutual inductance variation is estimated after extracting the sum of the faulty current pairs and applying equation (10). Fig. 5 demonstrates the dynamic degradation process starting from 1 %, and then experience some step changes to 2% and 10%. Again, the close agreement observed between the reference maximum mutual inductance variation rate to the estimated value, wherein the maximum error is only around 2%, successfully verified the effectiveness of the proposed quantitative electrical model and the \u201csoftware-based notch filtering\u201d technique. As shown in Fig. 6, an experimental setup is established with an 1-hp induction machine with an air gap length of 0.28 mm and two 6022-ZZ bearings mounted on the load side and the opposite side respectively. The bearing fault on the load side 6022-ZZ is created by either contaminating the bearing with powders consisting of tiny particles [17] or through electrolytic corrosion, while the opposite side bearing is kept healthy. To validate the proposed bearing fault quantification methodology via air gap displacement, two accelerometers are \u2022 installed close to the bearings on both sides of the motor caps to measured the real-time vibration signals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003753_s42835-020-00541-3-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003753_s42835-020-00541-3-Figure3-1.png", + "caption": "Fig. 3 Illustrating of no-load leakage flux", + "texts": [ + " According to divided areas, total air-gap leakage permeability can be calculated as, Then, main magnetic conductivity can be calculated as, where, bm is the width of PM, kg is the carter factor of closed slot, which is quoted by [18]. Hence, the air-gap leakage flux coefficient can be calculated as, At a certain position of rotor, a part of flux generated by one PM does not pass through the stator tooth and stator yoke to return to other PM not linking the windings in slots, which is zigzag leakage flux. Due to small difference between the number of poles and slots for SPMSM with similar number of poles and slots, there is not only air-gap leakage flux, but also bigger zigzag leakage flux. Figure\u00a03 shows the twodimensional finite element simulation calculation diagram of the SPMSM. It can be seen from the distribution of magnetic field that the zigzag leakage flux is much larger than the air-gap leakage flux, indicating the importance of zigzag leakage flux in this manuscript. The zigzag leakage flux of each tooth in one cell motor is different, including a large proportion of zigzag leakage flux and little air-gap leakage flux. So the zigzag leakage flux will not participate in electromagnetic energy conversion and the no-load back EMF will be weakened" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000671_012029-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000671_012029-Figure3-1.png", + "caption": "Figure 3. Contact points and angles (C0& \u03b10=0 \u0219i C1& \u03b11) for first two contacts, 0 and 1.", + "texts": [ + " (7) By transforming xjyj in xy( xyXjYjT ) result ii ii xyXjYj cossin sincos T . (8) Also, the relation (6) and (7) led to ijAijAA ijAijAA Cj Cj cossinrsincosrR sinsinrcoscosrR y x . (9) The previous mathematical model is applied to a real case and led to results presented in Table 1. 3.2. Method II The previous mathematical model for determining the contact angle (\u03b1) and also the contact points coordinates (xcj, ycj), can be verified using the second method. In Figure 3, the chain link and sprocket ensemble is considered rotated counter clockwise with angle \u03b8 o relative to the mounting position; so, the first position is for A0B0C0 and \u03b10 when the angle \u03b8 o is imposed. PRASIC IOP Conf. Series: Materials Science and Engineering 514 (2019) 012029 IOP Publishing doi:10.1088/1757-899X/514/1/012029 Following the Figure 3, the center coordinates for curve radius (A0,A1) and bush center (B0, B1) can be written as sinRy cosRx :A AA AA 0 0 0 ; sinrrRy cosrrRx :B BAAB BAAB 0 0 0 , (10) 1 1 1 1 1 sinRy cosRx :A AA AA ; 101 101 1 sinlyy coslxx :B BB BB . (11) Using the right line ecuation \u03941 for point C1 result 11 11 11 11 AB AB AC AC yy xx yy xx . (12) Using the circle ecuations for radius rA and rB, for point C1, result 22 11 2 11 AACAC ryyxx , (13) 22 11 2 11 BBCBC ryyxx . (14) Also, from relation (13) and (14) result 22 1111111111 22 ABBACBABACBA rryyyyyxxxxx ", + " Analyzing the Figures 4, 5 and 6 can be observed that: the contact angle increases linear with the increasing of the sprocket rotation angle; the contact point coordinates on x axis decreases with the increasing of the sprocket rotation angle; the accentuate decrease is at contacts considered in increasing order; the contact point coordinates on y axis increases with the increasing of the sprocket rotation angle; the accentuate increase is at contacts considered in increasing order. PRASIC IOP Conf. Series: Materials Science and Engineering 514 (2019) 012029 IOP Publishing doi:10.1088/1757-899X/514/1/012029 For results comparison, in Table 6 are presented the point coordinates and also contact angles values. It is considered that, the first contact starts from initial position on the vertical axis, \u03b10 = 0, (Figure 1) and vary along the sprocket angular pitch, \u03c4 (Figure 3). PRASIC IOP Conf. Series: Materials Science and Engineering 514 (2019) 012029 IOP Publishing doi:10.1088/1757-899X/514/1/012029 The correct functioning of a bush chain transmission is given, among other functional parameters, by the contact point / angle between the bushing and the sprocket, knowing that with the increase of the contact angle the friction surface between them increases. The design of a chain drive with bushings must also take into account the contact angle or contact point between the bushing and the sprocket as they influence the transmission dynamics (vibrations and wear) and can compensate for deviations induced by assembly or manufacturing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002413_012105-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002413_012105-Figure4-1.png", + "caption": "Fig. 4 enlarged drawing of maximum working stress position of robot mechanical structure (wind power level 12, wind pressure 1500 Pa)", + "texts": [], + "surrounding_texts": [ + "In this paper, the finite element software ANSYS was used to analyze the static characteristics of the mechanical structure of a transmission line inspection robot, and to investigate whether the static strength of the structure under the air wind load meets the design requirements. The main conclusions are as follows. (1) Under the action of wind load with different wind pressure, the maximum working stress of the robot is located near the left wheel. (2) With the increase of wind power level, that is, the wind pressure, the maximum working stress value of the mechanical structure of the robot is also increasing. When the wind power level reaches 12, that is, the wind pressure increases to 1500 Pa, the maximum working stress value of the structure also reaches the peak value, which is 209.854 MPa, less than the allowable stress of the material, and the static strength of the material meets the use requirements." + ] + }, + { + "image_filename": "designv11_80_0002469_issi47111.2019.9043711-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002469_issi47111.2019.9043711-Figure1-1.png", + "caption": "Fig. 1. Body-fixed frame and earth-fixed frame for the quad-rotor UAV.", + "texts": [ + " And then a fixed-time controller is constructed with less conservativeness by utilizing the faults estimation information from the observer. Notation 1. In this brief, \u2016 \u00b7 \u2016 is utilized for the Euclidean norm of vectors and matrices. For any x = [x1, x2, x3] T \u2208 R 3, we define x\u03b1 = [x\u03b11 , x\u03b12 , x\u03b13 ] T , sgn(x) = [sgn(x1), sgn(x2), sgn(x3)] T , and sig\u03b1(x) = [|x1| \u03b1sgn(x1), |x2| \u03b1sgn(x2), |x3| \u03b1sgn(x3)] T , where \u03b1 \u2208 R and sgn(\u00b7) denotes the sign function. A quadrotor UAV model is depicted in Fig. 1, which include a rigid cross frame and four rotors. For the purpose of describing the dynamics and kinematics, two orthogonal coordinate systems are introduced, the body-fixed frame is represented by OXBYBZB and the inertia frame is represented by OXEYEZE which is fixed to the earth. To describe the quadrotor attitude dynamics, various methods of description have been proposed, such as Unit Quaternion, Modified Rodriguez Parameters (MRP) and Euler Angles, etc. One of the most used kind of these methods is Euler Angles, and it is described through the Euler angles roll, pitch and yaw (\u03c6, \u03b8, \u03c8)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure46.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure46.1-1.png", + "caption": "Fig. 46.1 a Geometry of the serrated cutter with run-out. Figure is adapted and modified from [17]. b Cross-sectional view at height z with run-out", + "texts": [ + " This is followed by the main conclusions in Sect. 46.4. This section describes the geometry of serrated cutters considering the influence of radial run-out. Models described in this section build on the classical work done by Merdol and Altintas [11], and by Dombovari et al. [12], and also on our own earlier reported work [17], by including the influences of run-out. 46.2.1 Modelling of Serration Profile A schematic and a cross-sectional view of a sinusoidal serrated cutter with run-out is shown in Fig. 46.1. Due to the run-out, cutter geometrical centre, denoted by Og, deviates from spindle rotation centre Or, by a constant radial deviation, q:. The run-out angle between direction of radial offset (deviation) and the nearest tooth at the bottom of the tool is denoted by d: We assume that there is no axial run-out or any cutter geometric axis tilt; i.e. q is constant along the cutter height. Due to the run-out, since the whole system rotates about the spindle axis through Or; it is convenient to model all things with respect to Or, and hence the xyz coordinate frame is attached to Or as shown. The cutter can have N number of flutes (teeth), but as an example, only three (ith, i\u00fe 1\u00f0 \u00deth, and i\u00fe l\u00f0 \u00deth) flutes are shown in Fig. 46.1. As there is a wavy surface along the flute of the serrated cutter, the local geometrical radius changes along the flute and the height. The local geometrical radius which is measured with respect to geometrical axis (dotted line through Og shown in Fig. 46.1a) which is parallel to zaxis for the ith flute at the height z is defined as: 46 Investigations on the Influence of Radial Run-Out on Cutting \u2026 533 Rg i \u00f0z\u00de \u00bc D 2 DRg i \u00f0z\u00de \u00f046:1\u00de where D is shank diameter (nominal diameter) of the cutter, and DRg i \u00f0z\u00de is the variation in local geometrical radius for the sinusoidal serration profile: DRg i z\u00f0 \u00de \u00bc A 2 A 2 sin 2pz k cos g wi \u00fe p 2 \u00f046:2\u00de where A is the serration amplitude, which is half of the peak-to-peak serration height; k is the wavelength; g is the helix angle; and wi is the phase shift due of the serrations on different flutes, explained in [17]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001748_s1052618819060074-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001748_s1052618819060074-Figure4-1.png", + "caption": "Fig. 4. Theoretical working area of the 2-PUS.2-P(US)2 mechanism under the horizontal orientation of the output link as compared with its overall dimensions.", + "texts": [ + " Let us note that, using this approach, there is a theoretical possibility of missing the points of special positions if the pitch of the partition grid (exhaustive search) of the working zone is too large. The coincidence of the sign of the determinant in two adjacent points, theoretically, does not guarantee that a special position between them is absent. If the grid pitch is sufficiently small, then the probability of missing a special position is also low. By analyzing all points of the working area, one can make a conclusion on the influence of special positions on its size and shape [16]. Let us consider the 2-PUS.2-P(US)2 mechanism (Fig. 4), which has two simple (i = 2 and i = 3) chains and two double (i = 1 and i = 4) chains. The chains are arranged so that the rotation of the output link is possible merely around axis y'. The geometric dimensions (m) are the following: the lengths of links (they \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 = = \u2202 \u2202 \u2202 \u2202 \u2202 \u2202 \u2026 \u2026 1 1 1 1 1 1 0 0 \u03c7 \u03c7 , 0 0 , 0 0 \u03c7 \u03c7 n n n n n n F F F q F F F q A B JOURNAL OF MACHINERY MANUFACTURE AND RELIABILITY Vol. 48 No. 6 2019 are identical for all chains) are the following: lAB = 0", + "04, yB3 = 0.19, xA4 = \u20130.2, yA4 = 0, xB4 = \u20130.15, yB4 = 0; and the coordinates of points \u0421i in the initial position in the system Ex'y'z' are = 0.08, = 0, = 0.04, = 0.04, = \u20130.04, = 0.04, = \u20130.08, = 0, for all chains = lDE = 0.05. Here, the conditional points B and C are located in the middle of links B1B2 and C1C2 respectively, i.e., on the symmetry axis of the parallelogram of the double chain. To determine the theoretical working zone of the mechanism at zero rotation angle of the output link (Fig. 4), an iterative analysis in the range of coordinates was carried out: x = [\u20130.3; 0.3], y = [\u20130.3; 0.3], z = [\u20130.025; 0.5]. The grid pitch is 0.01 m along coordinates x, y and 0.025 m along coordinate z. Thus, in the volume limited by the limits mentioned, 81 862 points are analyzed. For investigation of the specific positions of types (1) and (2), the iterative analysis with the exhaustive search parameters mentioned above was carried out. Besides the horizontal position of the output link, four other examples with different rotation angles of the output link around axis y' were considered: 15\u00b0, 30\u00b0, 45\u00b0, and 60\u00b0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002215_s00170-019-04806-8-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002215_s00170-019-04806-8-Figure6-1.png", + "caption": "Fig. 6 Projections of the velocities of the contact points during the rotation around the hole\u2019s axis onto axes \u041e1\u0435 and \u041e1h", + "texts": [ + " 5) taking values of sin\u03c4 \u00bc b 0:5D and cos\u03c4 \u00bc S1 0:5D into account shall be transformed to the following form: V\u03c8 B1x \u00bc \u2212 S1sin\u03c8\u2212bcos\u03c8\u00f0 \u00de\u03c8 ; V\u03c8 B2x \u00bc \u2212 S1sin\u03c8\u00fe bcos\u03c8\u00f0 \u00de\u03c8 ; V\u03c8 Kx \u00bc 0:5Dsin\u03c8\u03c8 ;V\u03c8 B1y \u00bc S1cos\u03c8\u00fe bsin\u03c8\u00f0 \u00de\u03c8 ; V\u03c8 B2y \u00bc S1cos\u03c8\u2212bsin\u03c8\u00f0 \u00de\u03c8 ;V\u03c8 Ky \u00bc \u22120:5Dcos\u03c8\u03c8 ; V\u03c8 B1z \u00bc 0;V\u03c8 B2z \u00bc 0;V\u03c8 Kz \u00bc 0: \u00f013\u00de To define projections of velocities V\u03c8 B1 and V \u03c8 B2 onto moving axes of coordinates\u041e2\u03b5 and\u041e2\u03b7, they are first divided into two components. One of these components is parallel to axis \u041e1\u0435, and the second one is parallel to axis \u041e1h (Fig. 6). V \u03c8 B1 \u00bc V \u03c8 B1\u0435 \u00fe V \u03c8 B1h; V \u03c8 B2 \u00bc V \u03c8 B2e \u00fe V \u03c8 B2h: The values of these components define the projections of velocities to auxiliary axes \u041e2\u0435 and \u041e2h (see Fig. 6). V\u03c8 B1e \u00bc \u2212V\u03c8 B1cos\u03c4 \u00bc \u22120:5D\u03c8\u02d9 O1B 0:5D \u00bc \u2212S1\u03c8\u02d9 ; V\u03c8 B2\u0435 \u00bc \u2212V\u03c8 B2cos\u03c4 \u00bc \u22120:5D\u03c8\u02d9 O1B 0:5D \u00bc \u2212S1\u03c8\u02d9 ; V\u03c8 B1h \u00bc V\u03c8 B1sin\u03c4 \u00bc 0:5D\u03c8\u02d9 BB1 0:5D \u00bc b\u03c8\u02d9 ; V\u03c8 B2h \u00bc \u2212V\u03c8 B2sin\u03c4 \u00bc \u22120:5D\u03c8\u02d9 BB2 0:5D \u00bc \u2212b\u03c8\u02d9 : \u00f014\u00de Components V\u03c8 B1h and V\u03c8 B2h, in its turn, shall be decomposed into two components, one of which is parallel to peg \u041e2\u03b6 axis, the second is parallel to axis \u041e2\u03b7 V \u03c8 B1h \u00bc V \u03c8 B1\u03b6 \u00fe V \u03c8 B1\u03b7; V \u03c8 B2h \u00bc V \u03c8 B2\u03b6 \u00fe V \u03c8 B2\u03b7: Thus, velocitiesV\u03c8 B1 and V \u03c8 B2 of points\u04121 and \u04122 of rotation peg motion around hole \u041e1z axis shall be decomposed into three components. V \u03c8 B1 \u00bc V \u03c8 B1\u03b5 \u00fe V \u03c8 B1\u03b6 \u00fe V \u03c8 B1\u03b7; V \u03c8 B2 \u00bc V \u03c8 B2\u03b5 \u00fe V \u03c8 B2\u03b6 \u00fe V \u03c8 B2\u03b7: Projections of velocities V\u03c8 B1, V \u03c8 B2, and V \u03c8 K to moving axes of coordinates \u041e1\u03b5, \u041e1\u03b7, and \u041e1\u03b6 shall be obtained through adding projections of their components (see Fig. 6). V\u03c8 B1\u03b5 \u00bc V\u03c8 B1e \u00bc \u2212S1\u03c8\u02d9 ; V\u03c8 B2\u03b5 \u00bc V\u03c8 B2e \u00bc \u2212S1\u03c8\u02d9 ; V\u03c8 K\u03b5 \u00bc 0:5D\u03c8\u02d9 ; V\u03c8 B1\u03b7 \u00bc V\u03c8 B1hcos\u03b3 \u00bc b\u03c8\u02d9 cos\u03b3; V\u03c8 B2\u03b7 \u00bc V\u03c8 B2hcos\u03b3 \u00bc \u2212b\u03c8\u02d9 cos\u03b3; V\u03c8 K\u03b7 \u00bc 0; V\u03c8 B1\u03b6 \u00bc \u2212V\u03c8 B1hsin\u03b3 \u00bc \u2212b\u03c8\u02d9 sin\u03b3; V\u03c8 B2\u03b6 \u00bc V\u03c8 B2hsin\u03b3 \u00bc b\u03c8\u02d9 sin\u03b3: V\u03c8 K\u03b6 \u00bc 0: \u00f015\u00de Rotation of peg about its axis occurs with angular velocity \u03c6\u0307 \u00bc d\u03c6 dt . Velocities V\u03c6 K , V \u03c6 B1, V \u03c6 B2 of contact points \u041a, \u04121 and \u04122 while moving are located in the plane of aligned peg end \u041e2\u03b5\u03b7, and are equal in magnitude: V\u03c6 K \u00bc V\u03c6 B1 \u00bc V\u03c6 B2 \u00bc 0:5d\u03c6\u02d9 : Projections of these velocities to peg \u041e2\u03b6 axis are equal to zero: V\u03c6 B1\u03b6 \u00bc V\u03c6 B2\u03b6 \u00bc V\u03c6 K\u03b6 \u00bc 0: Projections to moving axes \u041e2\u03b5 and \u041e2\u03b7 (see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002272_ifeec47410.2019.9014659-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002272_ifeec47410.2019.9014659-Figure3-1.png", + "caption": "Fig. 3. Illustration of baseline PM machine: (a) stator/rotor topology, and (b) windings configuration.", + "texts": [ + " It is important that secondary harmonic magnitudes are dependent on load conditions because of flux saturation. In the next section, the influence of secondary inductance harmonics on the position estimation is investigated based on FEA. By identifying all infeasible estimation regions, an optimal current control is developed to maintain the saliency-based drive for full load operation. III. FEA SIMULATIOM A 5 kW IPM machine model is built for FEA simulation in this section. Key machine specifications are listed in Table I. The geometric feature of the machine is illustrated in Fig. 3. This machine contains an 8-pole IPM rotor with V-shape topology and a 12-slots stator with concentrated windings. Fig. 4 illustrates (a) \u03b1-\u03b2 frame and (b) d-q frame simulated inductance waveforms at no load. In this simulation, the flux saturation effect is considered based on FEA. Ideally as mentioned in (2), L\u03b1\u03b1 and L\u03b2\u03b2 inductance should have sinusoidal waveform without saturation. However, visible waveform distortions are observed in Fig. 4(a). More importantly in (b), Ldd and Lqq inductance all contain periodic harmonics and mutual-inductance Lqd appears" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000613_1.5112677-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000613_1.5112677-Figure1-1.png", + "caption": "FIGURE 1. CAD-Modell of the UAV (a) and details of the bonnet (cowling) and the internal structure (stringers) (b)", + "texts": [ + " Furthermore, it has been shown that the material is very suitable for post-processing, as it is easy to rework (e.g., by grinding or bonding). So far, however, BJ has been used only for relatively small tools. Therefore, in this contribution it will be investigated whether the BJ process is also suitable for the additive manufacturing of larger tools for the thermoforming of plastic sheets. 150001-2 As an example of this case study was used the bonnet (cowling) of an Unmanned Aerial Vehicle (UAV). This UAV is being developed as part of the international research project ELCOD (see Fig. 1). The aim of this project is to design and to test a UAV that can remain in the air for a long time, e.g. to carry out a measurement of the pollutant load. The UAV has a wingspan of approx. 2.56 m and a weight of only 14 kg. A payload bay of 250x300x200mm is provided for measurement equipment. The structure and planking of the UAV is largely made of wood, as this material is light and flexible while providing sufficient strength. For the manufacturing of the cowling, however, wood is not suitable, because it is difficult to bring to the complex shape at the top of the UAV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001497_sled.2019.8896303-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001497_sled.2019.8896303-Figure4-1.png", + "caption": "Fig. 4. Working principle of the magnetic levitation setup including acting forces.", + "texts": [ + " Nevertheless, no assumptions on the current ripple shape or restrictions on the current dynamic have been made, therefore allowing an estimation of the inductance with more precision and higher bandwidth. Moreover, this technique allows to identify and therefore compensate for the resistive component, which can change during operation due to self-heating or due to an induced Back-EMF during movement. The working principle of the electromagnetic levitation setup along with the acting forces is shown in Fig. 4. It consists of a coil with a ferrite core and an object to be levitated, which is in this case a hollow steel ball. The electromagnetic levitation system can be divided into an electric subsystem, described by Eq. 2, and a mechanical subsystem, which can be modeled by: F = m \u00b7 x\u0308, (16) F = m \u00b7 g \u2212 Fm(x, i). (17) The reluctance force Fm between the coil and the ball is depending on the position x of the ball and the driving current i in the coil. Many works such as [3] consider a phenomenological model for this force with a quadratic dependency on the current and a inverse quadratic dependency on the position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003504_s00773-020-00763-0-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003504_s00773-020-00763-0-Figure5-1.png", + "caption": "Fig. 5 Composition of a typical two-stroke marine diesel engine for CMV modelling", + "texts": [ + " The essential prerequisite for state observer design is the choice of appropriate propulsion engine model providing mutual relationships between the variables of interest with sufficient accuracy. The cycle-mean value (CMV) modelling approach provides the engine cycle-averaged temporal evolution of the engine internal states neglecting their in-cycle variations. This is due to the assumption that air and fuel mass flows pass through the engine continuously irrespective of the engine 1 3 cycle intermittent nature. The CMV model is constructed decomposing the engine into a number of components that are described by generic and reconfigurable mathematical models. Figure\u00a05 illustrates the main components considered in the CMV model. These are the cylinder, the scavenging and exhaust receivers, the compressor and turbine of the turbocharger, and the scavenging air cooler. The engine bmep is calculated as a difference between the indicated mean effective pressure (imep), Pi, and friction mean effective pressure (fmep), Pf. In turn, the imep is considered proportional to fuel pump index, hp, and fmep is considered to be a linear function of fuel pump index, hp, and engine rotational speed, ne; hence, where \u03b7c is the combustion efficiency, Pi0 is the imep corresponding to the maximum continuous rating (MCR) of the engine, coefficients kf0, kf1 and kf2 are selected such that the fmep calculation is consistent with the experimental measurements, Ga is the air mass flow rate through the engine, and Gf is the fuel mass flow rate delivered to the engine by fuel pump" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001707_icems.2019.8921756-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001707_icems.2019.8921756-Figure4-1.png", + "caption": "Fig. 4 Two models of stator core", + "texts": [ + " Stator tooth structure is complicated and has a great influence on the natural frequency of the stator ,so it can not be ignored when calculating the natural frequency. In order to improve the calculation accuracy, the mass and stiffness of the stator teeth are considered. The principle that the stator yoke and the stator teeth are equal in mass and stiffness are equivalent to a coaxial ring, and the electromechanical analogy method is used to derive the analytical formula of the stator natural frequencies. 2D model and simplified model are shown in Fig.4. Fig.3 shows the two models of the stator core ,which are equal in mass and stiffness. The formula can be obtained from the simplified model using electromechanical analogy method. Electromechanical analogy is a method of solving mechanical problems by equating mechanical variables into circuit variables. Then the variables are obtained by the circuit principle, and finally the equivalent is the mechanical quantity. The vibration displacement is equivalent to the current of the branch in the circuit, and the current is taken to obtain the vibration displacement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002392_978-981-15-1293-3_8-Figure1.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002392_978-981-15-1293-3_8-Figure1.3-1.png", + "caption": "Fig. 1.3 Diagram of an atomic force microscope (AFM)", + "texts": [ + " It was the first time that a microscope instrument like the STM could provide the ability to image small molecules and individual atoms, and the STM is still commonly used, especially to study the physics of semiconductors and metals [5]. It is able to directly obtain atomic-resolution three-dimensional (3-D) images of solid surfaces [6]. Much of the STM operation is conducted in ultra-high vacuum (UHV) and at low temperatures. The major downside of the STM is the requirement of conductive samples, thus ruling out most of its possible applications in important fields [5]. The invention of the STM was followed a few years later by the born of the Atomic Force Microscope (AFM) or Scanning Force Microscope (SFM) (Fig.\u00a01.3). The AFM is the most popular type of SPM family because, different from the STM, it can be utilized with non-conductive samples, and so has extensive applicability. A drawback of the AFM is that the soft nature and stickiness of the biological samples can interfere with the tip, so atomic resolution could not be reached. A noteworthy improvement in the study of biological as well as other soft materials is made by the creation of tapping-mode AFM, in which the probe or the tip oscillates at a resonant frequency and at amplitude setpoint whilst scanning across sample surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure3-1.png", + "caption": "Fig. 3. Proposed Model", + "texts": [ + " In Figure 1, it observed that the magnet structure of the PMMs adopted the conventional model. In the model, the height of the magnet was homogenous in all parts of the magnet structure. Thus, the height in the edge of the magnet is the same as the height in the center. Figure 2 shows the one step slot employed in the magnet edge. By using one step slot causes the magnetic flux distribution to become changing and decreasing in the edge of the magnet. Also, the total magnetic flux flowing into the air gap becomes reduced. Figure 3 represents the magnet model of PMM proposed studied in this paper. In this model, a-two steps slot was employed in the magnet edge to achieve and reduce the total magnet flux flowing into the air-gap of the PMM. Using the two steps slot in the magnet edge might expect the reduction of the amplitude of the CT. All the CT of PMM models Authorized licensed use limited to: City, University of London. Downloaded on July 10,2020 at 12:04:30 UTC from IEEE Xplore. Restrictions apply. studied and compared in this paper. They are the structure of 48 slots /6 poles. Figure 3 represents the magnet model of PMM proposed studied in this paper. In this model, a-two steps slot was employed in the magnet edge to achieve and reduce the total magnet flux flowing into the air-gap of the PMM. Using the two steps slot in the magnet edge might expect the reduction of the amplitude of the CT. All the CT of PMM models studied and compared in this paper. They are the structure of 48 slots /6 poles. TABLE I. RESPONSE SURFACE METHOD (RSM) In the beginning, to achieve the CT reduction of the proposed model, the structure of the proposed model was designed using Response Surface Method (RSM)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003937_ecce44975.2020.9235661-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003937_ecce44975.2020.9235661-Figure5-1.png", + "caption": "Fig. 5. 3D FEA model. Fig. 6. 3D FEA model side view.", + "texts": [], + "surrounding_texts": [ + "The circuit parameters are not uniquely defined and depend on k. Thus, to match the physical conditions more closely, the static 3D FEA results are used to calculate the main inductance. The radial component of the magnetic flux density penetrating the rotor surface facing one pole of the machine is hereby defined as the main magnetic flux density. L\u00b5 is then calculated using the number of turns N : \u03a8\u00b5 = N \u222b ~Br \u00b7 d ~A (7) \u03a8\u00b5 Is,nl = L\u00b5 (8) The ratio k is then calculated using (6)." + ] + }, + { + "image_filename": "designv11_80_0000407_978-3-030-23672-4_9-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000407_978-3-030-23672-4_9-Figure2-1.png", + "caption": "Fig. 2. Fault simulator platform.", + "texts": [ + " For diagnosing rotating machinery, Spectra Quest\u2019s system is used to collect a large database of vibration signals of different faulty elements of the bearing. Different states of measures are presented in Table 1. U205 is the reference of the used roller bearing. Measurements were taken by piezoelectric sensor (ICP) with respectively the frequency and measurement ranges 0. 3 at 10000 Hz and \u00b1 50 g, its sensitivity is 100 mV/g and it is mounted on the bearing by stud. The directions of the sensors are illustrated in Fig. 2; the measurements are taken simultaneously with three sensors at the same time. The parameters for the Machine Fault Simulator (MFS) bearings and the calculated frequencies of the bearing: the Fundamental Train Frequency (FTF), The ball Pass Frequency Outer (BPFO), The ball Pass Frequency Inner (BPFI) and The ball Spin Frequency (BSF), (see Table 2): In order to illustrate the signals measurements, an example plots are shown in this section with 15 Hz of rotating speed. A comparison of the healthy signal and the combined faults state signal is illustrate in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002376_b978-0-08-102941-1.00017-1-Figure17.11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002376_b978-0-08-102941-1.00017-1-Figure17.11-1.png", + "caption": "Figure 17.11 A strut tensile sample (right), the arrangement of the five struts in the gauge length (middle), and their relationship to an octet truss unit cell (left).", + "texts": [ + " As noted by previous authors, additively manufactured lattice structures are subject to a wide variety of imperfections, including under- or oversizing of strut diameters, nonuniform strut cross-sectional areas, waviness along the length of each strut, surface roughness (especially on horizontal struts), and porosity [26e29], which can have size-dependent effects on the mechanical properties of the lattice structure. To investigate this effect in depth, the mechanical properties of the struts were investigated experimentally. To characterize the properties of the struts, including all of their built-in imperfections, a strut tensile sample was designed, as shown in Fig. 17.11. Multiple struts were 3 In Fig. 17.10, it is evident that the FEA underpredicts the displacement of the lattice structures. The explanation is due to imperfections in the experimental specimens. Unlike the CAD models on which the FE simulations are based, the DMLS lattice structures are not perfectly flat on the top surface. As a result, the compression platen engages the DMLS lattice structure gradually, as represented by the lower stiffness region near the origin. As the lattice structure is fully engaged, it transitions to a linear forceedisplacement regime", + " included in each strut tensile sample because they are less susceptible than individual struts to damage during fabrication and removal from the build plate. Five struts were arranged in a staggered, symmetric pattern so that dual digital image correlation (DIC) cameras could observe all struts during tension testing; additional struts would have obscured other struts from the view of the dual cameras. The struts were arranged in parallel, as opposed to the octet unit cell arrangement on the left in Fig. 17.11, because individual unit cells are subject to distorted boundary conditions when they are tested in isolation from a periodically repeating lattice structure [49]. The length of each strut was specified as 3.5 mm to match the length of a strut in a unit cell of 5 5 5 mm, as utilized in the previous lattice structure experiments, and the diameter was specified as 0.82 mm, similar to the 0.75 mm value specified for the lattice structure experiments, to correspond to a lattice structure of approximately 25% relative density" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003929_j.pnucene.2020.103531-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003929_j.pnucene.2020.103531-Figure3-1.png", + "caption": "Fig. 3. The (a) schematic diagram and (b) the real soft fingers mold.", + "texts": [ + " Afterwards, the prepared silicone/W-based composites samples were ready for use in the fabrication of the soft fingers tested in this paper. The full details of the seven types of samples used in this study are shown in Table 2. The soft robots used in this study adopted the fast PneuNets (FPN) structural design (Mosadegh et al., 2014), with only the bottom part connected to the base. A schematic and an image of the real soft finger mold (with dimensions of 84 mm \u00d7 20 mm \u00d7 11.7 mm) used in this experiment are shown in Fig. 3. The system consisted of an extensible top layer, a middle layer, and an inextensible bottom layer. The gas chambers were connected through a single channel. The top and bottom layers were made of ductile silicone, and the middle layer was made of acrylonitrile butadiene styrene (ABS) films. The top layer was composed of five parts, one pedestal for fixing the internal gas chamber of the top layer and four side walls. The molds for the top and bottom of the soft fingers apparatus were made using 3D printing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000926_s11249-018-0989-y-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000926_s11249-018-0989-y-Figure3-1.png", + "caption": "Fig. 3 Air bearing surface of the slider", + "texts": [ + " Sealed connection boards and O-rings were used to prevent helium from leaking out of the chamber during the experiments. Commercially available aluminum disks of 3.5\u00a0in. (95\u00a0mm) diameter and 50\u00a0mils (1.27\u00a0mm) thickness were used in the experiments. The disk was coated with an 18\u00a0\u00c5 thick diamond-like carbon (DLC) protective layer and a 12\u00a0\u00c5 thin perfluoropolyether demnum lubricant (Daikin Industries, Ltd). The bonded lubricant ratio of the disk was about 80%. The slider surface was coated with a protective carbon overcoat (DLC) of approximately 2\u00a0nm. In Fig.\u00a03, the schematic of the air bearing surface of the slider is shown. We denote the leading edge of the slider by LE, the trailing edge by TE, and OE and IE denote the outer and inner edge of the slider, respectively. The nominal pitch and roll static angle of the sliders used in the experiments were 0.9\u00b0 and 0\u00b0, respectively. The same air bearing contour was used for both air and helium. In our study, a load\u2013unload cycle consists of the following steps: first, the disk was spun up to 5400\u00a0rpm. Thereafter, the slider was moved from the ramp to the outer diameter of the disk at a speed of 3\u00a0in" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003635_052061-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003635_052061-Figure6-1.png", + "caption": "Figure 6. Parameterization of the rolling element-raceway contact [4]", + "texts": [ + " To meet the static equilibrium contact angles rise due to displacements, deformations and misalignments of the rings caused by bending moments and axial loads [2, 7]. In this contribution the contact points are calculated by solving the static equilibrium of the rolling element raceway contact via finite element method. The approach is based on simplifications, such as the race control hypothesis that assumes no relative motion between the rolling elements and the raceway surface [13]. The model further addresses highly loaded blade bearings at low oscillation speed. Therefore, dynamic accelerations due to centrifugal and inertial effects are neglected. Figure 6 shows the geometry of the rolling element-raceway interaction used in this approach, generally described by Gupta in \u201cAdvanced Dynamics of Rolling Elements\u201d [4]. To determine the contact vectors the dynamic model described by Gupta was adjusted to the shapes of the raceways of four-point contact ball bearings. Furthermore, it has been assumed that the inertial and the raceway coordinate systems have the same position. Due to the overlapping load zones of blade bearings (Figure 5), the resulting rolling element speed is calculated by vector addition of both diagonal The Science of Making Torque from Wind (TORQUE 2020) Journal of Physics: Conference Series 1618 (2020) 052061 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003616_j.matpr.2020.08.536-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003616_j.matpr.2020.08.536-Figure7-1.png", + "caption": "Fig. 7. Analysis Setting.", + "texts": [ + " Simulation and working animation are the main and important aspects of the Solidworks. Solidworks made the engineering easy which helps to Table 5 Worm and Wheel Analysis. Analysis Type Static Structural Meshing Method Beam Method Element Type Triangular Element Size & Type 2.45 mm & 3D Number of Elements 27,983 Solver Sparse Direct Number of Nodes 42,662 develop a system before it is being manufactured. We can resolve the errors which are made during the design phase which avoids defects in the components (Fig. 7). The component is studied before it gets manufactured which reduced the occurrence of errors made. Flow simulation can also be performed and can get the results even there are flow simulation features which can be selected prior to the simulation. It also shows the results displayed on the screen (Fig. 8). An effective analysis to our Gear arrangement and metal casing to withstand all sorts of it takes, we have used Ansys workbench 16.0 we have done the static structural analysis by fixing the spindle and given the dynamic load at track arm with the Triangular mesh" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001228_icmtma.2019.00164-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001228_icmtma.2019.00164-Figure1-1.png", + "caption": "Fig. 1. Dimensions of the designed ESPAR antenna.", + "texts": [ + " In this letter, we demonstrate how single-anchor indoor localization concept can be advanced in terms of achievable accuracy by using a base station equipped with the designed ESPAR antenna and by employing both, directional main beam and narrow minimum, radiation patterns. Measurements of the prototype performed in a real-world scenario indicate that even for the proposed simple localization algorithm based on signal strength recorded in the BS, one can easily obtain levels of accuracy similar to the one for the most sophisticated algorithm presented in [1]. The proposed design comprises 12 passive elements ESPAR antenna (Fig. 1) with one active monopole in the center of the 1536-1225 \u00a9 2015 IEEE. Translations and content mining are permitted for academic research only. Personal use is also permitted, but republication/ redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. ground plane being a top layer of the printed circuit board (PCB) base. The active element is fed by an SMA connector, while the parasitic elements can be connected to the ground or opened by the single-pole, single-throw (SPST) switches connected to the end of each of them at the bottom layer of the structure. Parasitic elements connected to the ground are referred to as reflectors because they reflect energy, while the opened elements are directors as the electromagnetic wave can pass through them. All switches are controlled by an external microcontroller, hence the actual configuration of the antenna can be denoted by the steering vector , where denotes the state of each parasitic element in Fig. 1: for th parasitic element connected to the ground and 1 for opened. The antenna was designed and simulated in FEKO electromagnetic simulation software tool. The antenna design is based on those proposed in [4] and [5] and employs 1.55-mm-height FR4 laminate with top-layer metallization. In [4], the number of 12 parasitic elements was proven to be optimal with respect to the narrowest main lobe and the lowest backward radiation at the center frequency equal to 2.4 GHz, hence the same starting configuration was chosen, and then the antenna was optimized to obtain various radiation patterns (see Section II-B) and satisfactory input impedance matching for all parasitic elements configurations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001763_aeat-04-2019-0087-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001763_aeat-04-2019-0087-Figure9-1.png", + "caption": "Figure 9 Laboratory test setup (breadboard) for experimenting with the edge controller", + "texts": [ + " This selection was done for the following reasons: the extremely high performance to price ratio; high EM noise immunity; inherent resistance to ionizing radiation damage up to 50 krad(Si) (Leite et al., 2017); low power consumption; versatility; adequate, comfortable and free programming environment; and accessible and inexpensive function-rich programmer and debugger. Experimental testing setups and results A few prototype boards of the edge controller v8.0 were built for the purpose of testing in laboratory environments. A test laboratory setup has been constructed on a plywood board with dimensions of 50 50cm. It is shown inFigure 9.This setup uses: an edge controller (in the center of the setup); 390 KV, 200 W motor turning a carbon fiber pusher propeller and the motor and propeller are enclosed in a protective cage; battery block, consisting of 6 Li-ion battery cells of the type 18650 (3400 mAh each); balanced battery charger; and master switch with a master current sensing shunt (Table I). The testing setup in Figure 9 has been used not only in a number of experiments and tests of the edge controller itself, but also of auxiliary devices, such as an innovative battery block for UAVs (Zabunov and Mardirossian, 2019), microcontroller connectivity over I2C interface, etc. The following table shows the results from an experiment where the edge controller drives the BLDCmotor mentioned above at different output voltages and hence different output power and RPM. The efficiency of the module is compared for two supply voltages corresponding to 2-cell and 3-cell Li-ion battery supply, namely, 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure44.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure44.5-1.png", + "caption": "Fig. 44.5 Case 1; a Mastercam generated tool paths; b scallop estimation using VERICUT; c scallop estimation using developed system; and d sidestep variation", + "texts": [ + " This was followed by extensive testing of the system on different kinds of surfaces produced with CNC programs with various toolpath strategies, like iso-parametric, iso-scallop, and adaptive planar [12]. The following case studies show the results and validation of the approach. The case study 1 refers to a relatively simple geometry with known values of scallops while case study 2 explores a complex bicubic Bezier surface with adaptive sidesteps generated using algorithm proposed by Dhanda and Pande [12] for three-axis CNC machining. In this case, the CNC part program was generated using Mastercam X5. The model has concave and convex portions (Fig. 44.5a). The scallop height was set to be 0.05 mm on a flat surface with sidestep 1.091 mm. It can go up to 0.1 mm at 45\u00b0 to the lay direction. 514 A. Kukreja and S. S. Pande The scallop option in the VERICUT X-Caliper was used to measure the scallop height formed by two parallel intersecting cylinders chosen manually. It then identifies the cylinders being measured by displaying intersecting circles on the machined part model [5]. Figure 44.5b shows the generated histogram of the values estimated by the VERICUT (Range: 0.0499\u20130.058 mm). The scallop heights were estimated on 160 points (15 planes) throughout the surface using the X-Caliper tool of VERICUT. The average scallop was found to be 0.0506 mm. It is a known fact that the concave region has larger scallop heights and convex region has smaller scallop heights than the flatter region for the same sidesteps. The values estimated using VERICUT does not show this trend. This is because it does not consider the curvature of the surface while estimating scallops. The measurement/estimation is very approximate with limited points and sometimes unable to cover the whole machined surface. Figure 44.5c shows the scallop height values estimated using the developed system (Range: 0.04\u20130.06 mm). The values below 0.05 mm indicate the convex regions and the values above 0.05 mm indicates concave regions. The scallop heights are calculated based on the sidesteps that are found using the method explained in Sect. 44.2.3. A similar trend is seen for the sidestep (Fig. 44.5d). In both the Fig. 44.5c and d, there is a spread but it is quite small. It is seen that the scallop frequency histogram from our system (Fig. 44.5c) has a well-defined symmetric peak around the expected scallop value (0.05 mm). The spread in the sidestep values is due to the approximation in generating STL model of the machined stock, and the slicing of the STL model which may induce 44 On Estimation of Scallop Height from CNC Part Programs 515 errors (in calculating local minima) as the plane (facet) and curve intersection was solved using numerical methods in the developed system. This error in sidestep values directly affects the scallop height estimation and therefore some values of scallop height (Fig. 44.5c) are outside the tolerance band (0.04\u20130.06 mm). In this case, a complex bicubic Bezier surface was designed (Fig. 44.6a) and a CNC part program was generated using an adaptive planar strategy [12]. The maximum permissible sidesteps in various regions were automatically calculated by the algorithm reported in [12] based on surface curvature while maintaining the surface finish (i.e., tolerance <0.05 mm). The sidesteps were found to be in the range of 0.75\u20131.5 mm to maintain the required surface quality [12]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002032_3352593.3352615-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002032_3352593.3352615-Figure4-1.png", + "caption": "Figure 4: Illustration of Wheel-Terrain contact optimization", + "texts": [ + " Our assumption here is that the contact position of the first wheel (C1) is given in rover reference coordinate frame R. Fig.3 shows the different frame attached with the 10 DOF rover and wireframe structure of rocker-bogie. The rover base configuration in vector form is denoted as \ud835\udc3b = [\ud835\udc4b, \ud835\udc4c, \ud835\udc4d, \ud835\udf19\ud835\udc65 , \ud835\udf19\ud835\udc66 , \ud835\udf19\ud835\udc67] \ud835\udc47 which is defined with respect to the global coordinate frame G, where [\ud835\udc4b, \ud835\udc4c, \ud835\udc4d] is the position of the origin of R, orientation angles [\ud835\udf19\ud835\udc65 , \ud835\udf19\ud835\udc66 , \ud835\udf19\ud835\udc67], heading angle \ud835\udf19\ud835\udc67, pitch \ud835\udf19\ud835\udc66, and roll \ud835\udf19\ud835\udc65 as shown in Fig. 4. Using the DH table the following transformation matrices can be obtained \ud835\udc47\ud835\udc37 \ud835\udc45; \ud835\udc47\ud835\udc35\ud835\udc56 \ud835\udc37 , \ud835\udc56 = 1,2; \ud835\udc47\ud835\udc46\ud835\udc57 \ud835\udc35\ud835\udc56 , j = 1,2,3,4; \ud835\udc47\ud835\udc46\ud835\udc58 \ud835\udc37 , k=5,6, where Bi denotes the bogie frames, D differential frame , and Sj, and Sk denote auxiliary steering frames (attached to rocker arm). The transformation from the coordinate frame of reference of rover R to the frame of reference of axle Ai of each wheel are calculated as \ud835\udc47\ud835\udc34\ud835\udc56 \ud835\udc45 = \ud835\udc47\ud835\udc37 \ud835\udc45\ud835\udc47\ud835\udc35\ud835\udc56 \ud835\udc37 \ud835\udc47\ud835\udc46\ud835\udc56 \ud835\udc35\ud835\udc56\ud835\udc47\ud835\udc34\ud835\udc56 \ud835\udc46\ud835\udc56 , \ud835\udc56 = 1, 2. (1) \ud835\udc47\ud835\udc34\ud835\udc56 \ud835\udc45 = \ud835\udc47\ud835\udc37 \ud835\udc45\ud835\udc47\ud835\udc35\ud835\udc56\u22122 \ud835\udc37 \ud835\udc47\ud835\udc46\ud835\udc56 \ud835\udc35\ud835\udc56\u22122\ud835\udc47\ud835\udc34\ud835\udc56 \ud835\udc46\ud835\udc56 , \ud835\udc56 = 3, 4. (2) \ud835\udc47\ud835\udc34\ud835\udc56 \ud835\udc45 = \ud835\udc47\ud835\udc37 \ud835\udc45\ud835\udc47\ud835\udc46\ud835\udc56 \ud835\udc37\ud835\udc47\ud835\udc34\ud835\udc56 \ud835\udc46\ud835\udc56 , \ud835\udc56 = 5, 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003112_j.mechmachtheory.2020.103903-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003112_j.mechmachtheory.2020.103903-Figure3-1.png", + "caption": "Fig. 3. Hybrid serial-parallel 5-DOF 2-coupled-Cartesian-manipulator with parallel-revolute-axes, rotated 0 \u25e6 around \u02c6 Z W .", + "texts": [ + " Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx proven below, since revolute R joint axes \u02c6 Y Bn \u2016 \u02c6 Y Dn are geometrically parallel, there is no parasitic-twist-angle \u03b8BD n Z between links L Bn and L Dn so a passive revolute R joint is not required between them. Therefore links L Bn , L Dn , L T are rigidly connected, forming a single rigid body as shown in drawing In Sm Figs. 3 D, 4 D. The 5-DOF hybrid serial-parallel 2-coupled-Cartesianmanipulator with parallel-revolute-axes of Figs. 3 , 4 have joint topology ( PPP U )(P PP U ) . Note that they have four revolute R joints (two R \u2019s per U ) compared to the non-parallel-revolute-axes configuration Fig. 2 , with a total of five revolute R joints. The manipulator in Fig. 4 is identical to the one in Fig. 3 except that it is rotated \u03b8A 2 Z = \u221290 \u25e6, \u03b8C 2 Z = \u221290 \u25e6 around the \u02c6 Z W axis. The link and coordinate symbols in Fig. 4 are labeled with subscript \u2018 2 \u2019 to distinguish them from the ones in Fig. 3 with subscript \u2018 1 \u2019. Fig. 5 fully parallel 4-DOF 4-coupled-Cartesian-manipulator . Common-link L T in Fig. 5 connects the two 2-coupledCartesian-manipulators from Figs. 3 , 4 . Together they form the 4-DOF fully parallel 4-coupled-Cartesian-manipulator of Fig. 5 with 2 T 2 R motion-type. The passive revolute R joint, along the common-link \u02c6 Z T axis, accommodates parasitic-twistangle \u03b8BD n Z between links L B 1 , L D 1 and L B 2 , L D 2 since the revolute R joint axes of the two manipulators from Figs", + " Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx 11 has only two prismatic joints since it is fixed along the vertical \u02c6 Z W axis. The intersecting-revolute-axes version in Fig. 5 B, 5D has joint topology (P P UR )3( PP P U) , which may be represented as follow, suggesting the name \u2018Xactuator\u2019: Controlling the prismatic joint linear position z W A 2 in Fig. 5 B, 5D along vertical link L W adds position control z W T along \u02c6 Z W of the common-link L T for 5-DOF in a hybrid serial-parallel configuration, with joint topology ( P P P UR )3( PP P U) . In Fig. 5 , the 2-coupled-Cartesian-manipulator from Fig. 3 interleaves with the one from Fig. 4 so that link L B 2 from the second 2-coupled-Cartesian-manipulator is between links L B 1 , L D 1 from the first 2-coupled-Cartesian-manipulator. Similarly, link L D 1 from the first 2-coupled-Cartesian-manipulator is between links L B 2 , L D 2 from the second 2-coupled-Cartesianmanipulator. This allows coaxial bearings R 1 , R 2 to be spaced far apart from each other, along the common-link L T , as seen in Fig. 5 A, to support high moment loads with high moment stiffness", + " Mathematical description Kinematic equations [55] provide basis for the design, simulation, control and analysis of the coupled-Cartesian- manipulator family. The coordinate frames of Fig. 7 illustrate the 2-coupled-Cartesian-manipulators of Figs. 2\u20134 that are composed of two limbs, { L An , L Bn , L T } and { L Cn , L Dn , L T } respectively. The nomenclature in Fig. 7 pertains to generic links L An , L Bn , L Cn , L Dn where n = { 1 , 2 } but it applies specifically to links L A 1 , L B 1 , L C 1 , L D 1 in Fig. 3 or links L A 2 , L B 2 , L C 2 , L D 2 in Fig. 4 . The two 2-coupled manipulators from Figs. 3 , 4 , joined together by a revolute R joint along the common-link L T longitudinal \u02c6 Z T axis, form the 4-coupled-Cartesian-manipulators in Figs. 5 , 6 , each with four limbs. The general mathematical equations, derived for the 2-coupled manipulators in Figs. 3 , 4 , also apply to the 4-coupled manipulators in Figs. 5 , 6 . Coordinate frames. The positions and orientations of the components of the manipulators are expressed relative to Cartesian coordinate frames as shown in Fig", + " Similarly, if common-link L T connects to coupler-link L Dn then the orientation angles \u03b8C n X , \u03b8D n Y of links L Cn , L Dn depend on orientation angle \u03b8C n Z of link L Cn , \u03b8Cn X = atan 2 ( sin ( \u03b8Cn Z ) x W \u0302 Z T \u2212 cos ( \u03b8Cn Z ) y W \u0302 Z T , z W \u0302 Z T ) , 0 \u2264 \u03b8Cn X \u2264 2 \u03c0 (16) \u03b8Dn Y = sin \u22121 ( cos ( \u03b8Cn Z ) x W \u0302 Z T + sin ( \u03b8Cn Z ) y W \u0302 Z T ) , \u2212 \u03c0/ 2 \u2264 \u03b8Dn Y \u2264 \u03c0/ 2 (17) For the coupled-Cartesian-manipulators analyzed here, angles \u03b8A n Z , \u03b8C n Z in Eqs. (14 - 17 ) are fixed. For example, \u03b8A 1 Z = 0 \u25e6, \u03b8C 1 Z = 0 \u25e6 in Fig. 3 and \u03b8A 2 Z = \u221290 \u25e6, \u03b8C 2 Z = \u221290 \u25e6 in Fig. 4 . Inverse kinematics, common-link L T position T W . Given common-link L T desired position T W , the positions A n W , C n W of links L An , L Cn are derived from Eqs. (12 , 13 ) A n W = T W \u2212 R W B n ZXY T Bn \u2212 R W A n ZX B n An (18) C n W = T W \u2212 R W D n ZXY T Dn \u2212 R W C n ZX D n Cn (19) Given desired common-link L T position T W and orientation \u02c6 Z W T , all of the terms on the right hand sides of Eqs. (18 , 19 ) are known from equations (5 , 6 , 14-17 ) and the fixed geometry of the links B n An , T Bn , D n Cn , T Dn " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002302_gncc42960.2018.9018642-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002302_gncc42960.2018.9018642-Figure2-1.png", + "caption": "FIGURE 2. The force analysis of the spiral drum.", + "texts": [ + " The drums connect the fuselage of shearer with two arms. The arms can adjust the height of the cutting drums by the hydraulic cylinder. The fuselage of the shearer rides on the scraper conveyor and moves along it. The coal exploitation can then be conducted with the procedure given above [19]. Fig. 1 shows the schematic diagram for the shearer. Because the cutting drums must bear many complicated stresses from the coal, analyzing the stress of the drums becomes more important. A. FORCE ANALYSIS OF DRUM The force of the cutting drums is shown in Fig. 2. Fig. 2 (a) and (b) represent the rear and front drums, respectively. In the figure, Xi, Yi and Zi are the lateral 2046 VOLUME 4, 2016 resistances traction direction resistance and cutting resistance of the i-th pick on the front drum, respectively. X \u2032 i , Y \u2032 i and Z \u2032 i express the resistances of the i-th pick on the rear drum. Rx , Ry and Rz are the resultant forces of all picks that are cutting the coal on the front drum, and R \u2032 x , R \u2032 y and R \u2032 z are the resultant forces on the rear drum. N1 and N2 are the numbers of picks of the front and rear drums, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001184_j.mechmachtheory.2019.103606-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001184_j.mechmachtheory.2019.103606-Figure11-1.png", + "caption": "Fig. 11. The 3-PSR parallel mechanism.", + "texts": [ + " The allowed movement is the rotation around the line formed by the two centers of the spherical joints. Thus, it is possible to replace the two spherical joints directly with one rotational pair. The spatial inverted triangle chain becomes to a PRS limb which everyone is very familiar with, as shown in Fig. 9 (c). The fifth combination ( C \u2207 C \u2207 C \u2207 ) would become the 3-PRS parallel manipulator as shown in Fig. 10 . Similarly, the spatial upright triangle chain can be equivalent to a PSR limb. The fourth combination ( C C C ) would become the 3-PSR parallel manipulator as shown in Fig. 11 . In the previous research, the established models of 3-PRS parallel mechanism are only suitable for the special structure (e.g., Fig. 10 (a)). For the more generic 3-PRS parallel manipulator shown in Fig. 10 (b), it is difficult to establish an analytical model because the coupling relationship between its output parameters has not been revealed. However, based on its equivalent mechanism (i.e. the fifth configuration), it is easy to deal with this problem. In addition, when it is necessary to solve calibration model of the linear Delta, 3-PRS and 3-PSR mechanisms, it can be equivalent to the mechanisms presented in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003452_0954405420951101-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003452_0954405420951101-Figure3-1.png", + "caption": "Figure 3. The coordinate transform between the rolling tools and rack-cutter.", + "texts": [ + " u1 is the ellipse parameter and the range of value is [u1min, p-u1min] with: u1min =p=2z1 tan (p=9) p=9\u00f0 \u00de \u00f02\u00de The ellipse is tangent to tooth flank and root surface, the radii a, b, d can be evaluated by: m+ c3m d= b r rf 3 cos (u1min)= b3 sin (u1min)+ d rf 3 sin (u1min)= a3 cos (u1min) 8< : \u00f03\u00de Based on equation (3), one can obtain a= rt 3 tan (u1min) b=((1+ c)3m+ rt 3 cos (u1min) r)= (1 sin (u1min)) d=(r rt 3 cos (u1min) (1+ c)3m3 sin (u1min)) =(1 sin (u1min)) 8>>>>< >>>: \u00f04\u00de The geometry of the gear rack-cutter is related to the proposed tooth profile. The coordinate transformation from the proposed gear to rack-cutter is shown in Figure 3. Coordinate systems S1 and Sc are rigidly connected to the gear and rack-cutter that perform rotational and translational motions with respect to the fixed coordinate system S. The clockwise rotation angle of the gear is noted as fcwhen the horizontal movement distance of rack can be represented as lc = r1fc. A point M in coordinate system S1 is represented by position vector O1M * =M1. The equation of surface in system S1 can be represented with two independent parameters as M1(u 1 1, u 1 2). The conjugate rack surface can be obtained by coordinate transformation based on the concept of the envelope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002768_00207721.2020.1723732-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002768_00207721.2020.1723732-Figure1-1.png", + "caption": "Figure 1. AUV coordinate system.", + "texts": [ + " In such discrete versions, the delays may have impact on the systems to be controlled, but this fact should be handled in terms of the sampling theory more than the approach proposed in this study. We agreed with the reviewer that taking into account delays in the righthand side of the system to be approximated is a really interesting problem, but it is outside the scope of the present study. In order to validate the performance of the RDE based on T\u2013S fuzzy control, an AUVmodel is presented, see Figure 1. This AUV is equipped with two propellers to control the vehicle in x\u2212y positions. In order to obtain a model of this AUV in the horizontal plane, two frames: the inertial reference frame XI \u2212 YI , and the body frame XB \u2212 YB are defined. The origin of the body frame is chosen as the centre of gravity of the AUV. In the inertial reference frame, XB can be regarded as longitudinal axis (from aft to fore), YB is regarded as transverse axis (starboard direction). A simplified dynamic model with respect to the body frame in the horizontal plane is d dt p (t) = q (t)w (t) + \u03c4p (t) + \u03be (t) d dt q (t) = \u2212p (t)w (t) d dt w (t) = \u03c4w (t) y (t) = [ q (t) w (t) ] (56) Here p, q are linear velocities inXB\u2212axis and YB\u2212axis, respectively, w is the angular velocity with respect to CG, \u03c6 is the yaw angle in the inertial reference frame XI \u2212 YI , \u03c4p = Fp + Fs is the total force along to XB\u2212axis, Fp is the port thrust, Fs is the starboard thrust" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002653_012008-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002653_012008-Figure1-1.png", + "caption": "Figure 1. Three-wheeled tadpole design", + "texts": [], + "surrounding_texts": [ + "Three-wheeled vehicle is that operates on three wheels. Three-wheeled vehicle has two configurations which are the tadpole configuration and the delta configuration. A tadpole configuration is a threewheeled vehicle which has two wheels upfront and one wheel at the back while the delta configuration has two wheels at the back and one wheel at the front. The dynamics system of a tadpole configuration three-wheeled vehicle could be very complicated. Few matters to be considered are the turning and tilting system, components such as: turning radius, center of gravity, wheel track, etc. This components are then can be measured and determined to ensure the ride safety of the vehicle. For a three-wheeled vehicle to maintain its stability, a tilting system is developed. Tilting was meant to lower the CG so the maximum lateral acceleration that could be taken by the vehicle is higher which results in better stability, and so is the safety. The modelling aims to represent the physics that is happening with the vehicle for future research in active steering and tilting assist for a three-wheeled vehicle." + ] + }, + { + "image_filename": "designv11_80_0002099_icar46387.2019.8981581-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002099_icar46387.2019.8981581-Figure7-1.png", + "caption": "Fig. 7: Properties and uncertainties of the assembly process between the structural-frame and the panel", + "texts": [ + " After a successful assembly operation, there must be no clearance (< 0.3mm) between the parts and the contact forces must be uniformly distributed. Figure 6 illustrates the desired contact between the parts as well as the position of the force sensor. All the forces applied to the structural-frame are measured related to the sensor\u2019s coordinate system Sxyz. Due to the poor accuracy of robots, and the uncertainties of the position and orientation of the panels, the robot\u2019s trajectory must be adjusted during assembling processes. Figure 7 presents two possible contact scenarios: when the first contact happens with the upper surface of the structural-frame, and the second when it happens with the bottom surface. Since the frame is not perfectly rigid, it can deform during the assembly. Numerical simulations were done to estimate the deformation both scenarios, supposing the same normal force uniformly distributed at each surface. Figure 7 shows the scales of deformation for both cases. It can be noticed that the structural-frame is much more flexible (\u223c 200 times) at its upper region than at its bottom. Owing to these pose uncertainties, and the variable stiffness of the handled object, we decided to implement a multi-surface admittance control approach to achieve the contact task. This approach allows us to modify the stiffness behavior of the controller according to the stiffness of the frame\u2019s surfaces, and thus, guarantees the stability at the first contact and the success of the task", + " Looking at the performances of the force/torque control using the OSAC approach, the results show that the response varies a lot from one scenario to another. When the first contact is happening at the part\u2019s bottom (Scenario No.1), the controller is not able to stabilize the contact force. Differently, when it is happening at the top (Scenario No.2), the system can converge to the setpoints. However, when all the surfaces are being in contact (force/torque errors near to zero), the system starts oscillating without stabilization. This behavior can be explained since the stiffness of the structural-frame is very different at its bottom and top (Fig. 7). Therefore, even if the constants are set empirically, it is hard to believe that one set of constants could ensure the same performance in both scenarios. In the other hand, when the MSAC is used, the results show that the torque set-points are correctly reached in almost the same time (80-95 sec) for both scenarios. It is possible due to the definition of two independent controller structures, with different mass-spring-damper constants by surface. In this case, the torque controller of the top surface is set to be stiffer and damped (K0U2 and D0U2 higher than K0U1 and D0U1 )" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003239_tmag.2020.3012200-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003239_tmag.2020.3012200-Figure1-1.png", + "caption": "Fig. 1. Three-degree-of-freedom linear-rotational resonant actuator.", + "texts": [ + " To realize all of the resonant motion, we propose a structure in which four-phase coils and permanent magnets are arranged symmetrically. We also propose an inverter topology to control the four-phase coils with six-switches. In addition, finite element analysis (FEA) is performed to confirm the characteristics of the proposed actuator. Finally, linear resonant motion in X-Y direction and rotational resonant motion are verified through dynamic simulation based on FEA results. A. Overview of Proposed Actuator Fig. 1 demonstrates the proposed three-DOF LRA with four-phase coils. The actuator consists of stator, coil, permanent magnet, and moving core. four-phase coils are connected by a star-connection and wound counterclockwise when viewed from the top. Magnets placed on a-b-c phase with 120o interval are fixed to the moving core. In addition, three magnets are magnetized in the positive direction of the Z-axis while the others are magnetized in the negative direction of the Z-axis. Since the mover is constrained only in the Z-direction, various resonant motion can be realized on the X-Y plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure78.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure78.3-1.png", + "caption": "Fig. 78.3 Schematic of skeleton force diagram of Fibrizer hub and hammer", + "texts": [ + " From these parameters r/h; H/h; moment of inertia (I) and C which is distance from the neutral axis to the point of interest of hammer. By which, bending moment (M) and maximum shear stress dmax \u00bc kt MC I were evaluated considering a constant factor (kt) value as 1.58. Load analysis of current Fibrizer hammer was evaluated by using speed of Fibrizer (N) = 750 rpm and motor power = 2 MW. By calculating the center of gravity (Y = (A1Y1 + A2Y2 + A3Y3 + A4Y4)/A1 + A2 + A3 + A4)), angular speed (x = 2pN/60); linear velocity (v = x Y) and Torque (T = 60P/2pN). Figure 78.3 indicates the forces which are acting on Fibrizer hub and hammer during the process. There were three major forces acting on the hub and hammers such as force created by the pin (Fp), centrifugal force (Fc), and force generated by cane being shredded (Fs). Centrifugal force (Fc) was calculated by mrx2, where m is mass of hammer and r is distance of gravity center from rotation axis. Loading of one hammer during normal operation was value of tangential force, i.e., force created by pin (Fp) during conventional operation of Fibrizer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002376_b978-0-08-102941-1.00017-1-Figure17.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002376_b978-0-08-102941-1.00017-1-Figure17.2-1.png", + "caption": "Figure 17.2 (a) A refined approach to lattice structure modeling that explicitly considers the effect of process-induced variation on the behavior of the lattice structure and its fundamental constituents. (b) An integrated computational modeling approach that models struts, lattice structures, and their imperfections concurrently. (c) An upscaling computational modeling approach that creates separate stochastic models of the mesostructure and then utilizes them in macroscale models. (d) The approach utilized in this research in which experimental characterization of struts generates effective strut-level properties that serve as input to the macroscale lattice structure models.", + "texts": [ + " However, the primary assumptions are that (1) the properties, as derived from standard tensile bars, are representative of the mechanical properties of the material when fabricated in the form of small struts and (2) the geometry of the struts and their arrangement into unit cells match as-designed specifications precisely. All of the process-induced variability is assumed to be captured in the bulk material properties of the standard tensile bars. This approach is implemented in Sections 17.3 and 17.4, where it is found to overestimate the properties of lattice structures by 30%e40%, which is in agreement with similar studies reported in the literature [40,41]. A refined approach to modeling lattice structures is illustrated in Fig. 17.2. In this approach, the design engineer assumes that the AM process is responsible for variability in not only the base material properties but also the effective mechanical properties of the struts themselves and the unit cells in which they are arranged (i.e., the mesostructure). In this paradigm, small changes in processing conditions can induce significant changes in effective mechanical properties via a wide variety of mediating factors. As shown in Fig. 17.2, those mediators include porosity, inhomogeneities, microstructural changes in the base material, and surface roughness, shape imperfections, size imperfections, and joint morphology in the mesostructure. These mediators, in turn, are affected by processing conditions such as scan pattern, laser power, powder morphology, thermal conditions during the build, postprocessing steps (heat treatment, HIP), part orientation, build density, and feature and part sizes. Given the wide variety of mediating factors, it becomes very challenging to predict the mechanical properties of a lattice structure accurately. There are at least three categories of approaches to address this challenge, as outlined in Figs. 17.2(bed). The first approach (Fig. 17.2(b)), discussed briefly in the introduction, relies on integrated computational modeling of individual struts assembled into a largescale lattice structure [40,41,44]. Mediating factors, such as shape and size imperfections, are characterized typically from CT scans of representative lattices; their statistical distributions are quantified; and those distributions serve as inputs to stochastic FEA of an expansive lattice structure comprised of a large number of \u201cimperfect\u201d struts. The second approach (Fig. 17.2(c)), also known as stochastic upscaling [45], relies on building models of the homogenized or effective mechanical properties of unit cells, which capture the variability in their performance. These statistical effective models are then used to determine the properties of the overall lattice structure. The deterministic version of this approach is used widely for computationally efficient analysis of lattice structures and honeycomb materials [46,47]. The benefit of this approach is its computational efficiency relative to integrated approaches because it is not necessary to model stochastic fine-scale features directly when evaluating the properties of a complex lattice structure; homogenized or effective models of the struts or unit cells are used, instead. The third approach (Fig. 17.2(d)), which is adopted in this research study, relies on building experimental models of the effective mechanical properties of struts and utilizing those empirical models as direct input to computational models of an overall lattice structure. This approach is more computationally efficient than other approaches because it does not require any simulation of fine-scale features. Significant experimentation is required, however, to build the empirical models of strut behavior. This approach is implemented in Sections 17.5 and 17.6, where it is found to estimate the mechanical stiffness of lattice structures with an order of magnitude less error than the standard approach depicted in Fig. 17.1. In summary, the focus of this research is to demonstrate the approach depicted in Fig. 17.2(d) and contrast it with the conventional approach illustrated in Fig. 17.1. The conventional approach is implemented in Sections 17.3 and 17.4. Specifically, Section 17.3 describes fabrication and testing of lattice structures and standard tensile bars. Section 17.4 reports the results of FEA of the lattice structures based on the bulk properties derived from the standard tensile bars and compares those lattice structure properties to the experimental results. The approach depicted in Fig. 17.2(d) is implemented in Sections 17.5 and 17.6, which describe experimental testing of strut-level tensile specimens and incorporation of the experimentally derived effective mechanical properties of the struts into a more refined FEA of the lattice structures. To investigate the mechanical properties of DMLS octet truss lattice structures, a test specimen was designed and fabricated. As illustrated in Fig. 17.3, the compression test specimen comprised 7 7 7 octet truss unit cells. Each individual strut was 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002846_012072-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002846_012072-Figure9-1.png", + "caption": "Figure 9. The consequence caused by deformation on tire tread when the plate is moved. (a) An angle is produced because of deformation x; (b) The effective radius becomes longer than the free rolling radius (re > ro).", + "texts": [ + " In this research, the tire characteristic modelling is done only from a slip which equals to zero until a slip at maximum friction force. Therefore, it can be assumed that the total deformation happened along contact patch lt and will be the same as the plate displacement x. The is a part of the tire tread affected by deformation (stretching). ICMER 2019 IOP Conf. Series: Materials Science and Engineering 788 (2020) 012072 IOP Publishing doi:10.1088/1757-899X/788/1/012072 The effect of tire tread deformation toward distance travelled by the wheel is shown in figure 9. From figure 9(a) it can be seen that when the plate is moved, an angle appears. This angle is also shown in figure 8(b). Thus, if the wheel rotates with an angle , it will travel like if it rotates with an angle + . Therefore, the distance travelled by the wheel will be further compared to the wheel with no deformation . This effect also can be explained by figure (b). Deformation S makes the effective radius (re) longer than the free rolling radius (r0). As a result, with the same angular velocity, the braked wheel will travel further than the non-braked wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure7.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure7.2-1.png", + "caption": "Figure 7.2 Graph of function K(s).", + "texts": [ + "1 The Kernel of the Generalized Model of Elastic Foundation (Base) 389 The limiting transitions with \ud835\udf00\u2192 0 is accomplished (carried out, implemented) according to formula (7.1.13) K(s) = k0\ud835\udeff(s). The kernel (7.1.25) can be used for describing the stress-strain state for soils that are only a little cohesive. The general form of the kernel of base with the plane strained state, and with \ud835\udf08 \u2260 0, \ud835\udf00 \u2260 0 can be obtained by integration (7.1.21) taking into account the condition at infinity K(s) = \u2212 \ud835\udf03\u0393 (3 2 \u2212 \ud835\udefc ) (s2 + \ud835\udf002)\u2212\ud835\udf08\u22152 \ud835\udf082\ud835\udf08\ud835\udf0b\ud835\udf08\u22121\u22152\u0393(\ud835\udefc + \ud835\udf08) (7.1.26) and using well known formula from the theory of gamma-functions \u0393(1 + z) = z \u0393(z). (7.1.27) Figure 7.2 shows the graphs of the kernel (7.1.26) with the different values of the parameters \ud835\udf08 and \ud835\udf00. Along with the kernels foundation for the spatial and planar (two-dimensional) problems considered in the Cartesian coordinate system, it is also of interest to obtain the corresponding expressions for the tasks, formulated in polar coordinates. As it is well known, the distance between two points in the plan with the polar coordinates (r, \ud835\udf11) and (r1, \ud835\udf111) is determined by the formula: 7.1 The Kernel of the Generalized Model of Elastic Foundation (Base) 391 Therefore, the relationship between the load on the base p(r1, \ud835\udf111) and displacement of its surface w(r, \ud835\udf11) can be represented in the form of: \u222b 2\ud835\udf0b 0 \u222b \u221e 0 K ( \u221a r2 + r2 1 \u2212 2rr1 cos(\ud835\udf11 \u2212 \ud835\udf111)) p(r1, \ud835\udf111)r1d\ud835\udf111dr1 = w(r, \ud835\udf11) Using the integral representation (7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003074_032070-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003074_032070-Figure1-1.png", + "caption": "Figure 1. (a) A misaligned journal bearing, (b) texturing parameters for a journal bearing.", + "texts": [ + " investigated the textured surface and journal misalignment on the performance of hydrodynamic journal bearing. However, When the journal misalignment and elastic deformation are considered, whether the optimized texturing parameters can improve the bearing performance has not been reported. In this research, an EHL model of partially textured journal bearing is developed and validated. Based on the model, the influence of journal misalignment and elastic deformation on the performance of partially textured bearings is analyzed. The geometry of a misaligned journal bearing is shown in Fig.1 (a). The film thickness in journal bearing considering journal misalignment can be expressed as 0 0 0 0 1 cos ' cos 2 z h c e e L (1) where c is radial clearance, 0e is the eccentricity at the bearing axial mid-plane, 0 is the attitude angle, 'e is the magnitude of the projected journal axis on the bearing mid-plane, is the misalignment angle between the line of centers and the rear center of the misaligned journal, L is the bearing width, is circumferential angle. The dimensionless film thickness considering journal misalignment can be described as 0 0 0 0 1 1 cos ' cos 2 y h L (2) Where 0 is eccentricity ratio, ' is the misalignment eccentricity ratio presented by [10] m max ' ' ' e D c (3) Where Dm is dimensionless degree of misalignment, max' is the maximum possible value of ' . As shown in Fig.1 (b), The geometric parameters of the dimple include radius rp, dimple depth hp, and area density Sp, 2 p p 1 2S \u03c0 r L L . The angle \u03b81 and \u03b82 are the angular position of texture. ESAET 2020 Journal of Physics: Conference Series 1549 (2020) 032070 IOP Publishing doi:10.1088/1742-6596/1549/3/032070 In the Winkler elastic foundation model [11], the deformable object is modeled as a set of spring elements. The elastic deformation can be therefore expressed as ' pt E (4) Where p is pressure, t is the thickness of bearing bush, 'E is a combined elastic modulus [11]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000827_1350650119866040-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000827_1350650119866040-Figure3-1.png", + "caption": "Figure 3. Finite element model and the results for the solution of multi-interfacial contact mechanics of axle bearings assembly.", + "texts": [ + " NJ3226 near the wheel and NJP3226 away from the wheel are, respectively, referred as \u2020, I\u2020 in Figure 2(b). Under given load and material properties, the real contact performance depends largely on clearance, interference fit of parts, and fit differences between them. These design parameters are controlled in a certain range from assembly manual and industry standards in Table 2. Location number of rollers and radial load distribution in this case are shown in Figure 2(c). Full numerical finite element model of axle bearings assembly considering multi-interface contact is established as shown in Figure 3, considering the internal radial and axial clearances of bearings, the interference fit between inner ring and journal as well as the clearance between outer ring and axle box housing. The model contains roughly three million of mesh grids with the cantilever beam structure. Multi-interface contact mechanics model is employed to study the contact load distribution and deformation of the entire system under actual working conditions, in which the operating radial load applied on the axle housing is 89 kN calculated by UIC510-5 (standard of International Union of Railways) and related parameters as shown in Table 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001549_s11223-019-00110-8-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001549_s11223-019-00110-8-Figure10-1.png", + "caption": "Fig. 10. Thermal stress diagram of the ring gear at: (a) 50 ; (b) 100 .", + "texts": [ + " Thermal strain and stress diagrams obtained by simulation are shown in Figs. 9 and 10. The maximum thermal strain always appears near the top of the ring gear tooth and the minimum appears near the root of the tooth. This is because thermal expansion of a part relates to its shape. The radius of curvature of the tooth top area is negative, and the radius of curvature of the tooth root is positive. The thermal strain reduces the gap between the movable tooth and the ring gear and the movable tooth may be jammed. As can be seen from Fig. 10, the maximum thermal stress is always near the root of the tooth, while the minimum is near the tooth\u2019s top. This is because the radius of curvature of the tooth root is negative, so stress concentrates at this place. Therefore, the tooth root is a dangerous place of the thermal stress damage during the meshing of the movable tooth and the ring gear at high speed and heavy load conditions. 4.1. Contact Meshing Stiffness of the Meshing Pair. In the drive, the meshing contact between the wave generator and movable tooth, movable tooth and ring gear can be reduced to the Hertz line contact" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001054_elan.201900109-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001054_elan.201900109-Figure3-1.png", + "caption": "Fig. 3. Schematic diagram showing the design of the multiple narrow-ring electrodes. Electrode arrays (EQ, INC, and DEC) consist of concentric ring electrodes with constant spacings of 75 \u03bcm. The widths of the ring electrodes were constant, at 50 \u03bcm, for EQ, increased from 50 \u03bcm to 100 \u03bcm in 10 \u03bcm increments from the center to outer periphery for INC, and decreased from 100 \u03bcm to 50 \u03bcm in 10 \u03bcm increments for DEC. DISC had a radius of 500 \u03bcm. Effective working electrode areas (mm2): DISC, 1.19; EQ, 1.94; INC, 2.06; DEC, 1.82.", + "texts": [ + " Numerical simulations were performed using a 2D axisymmetric geometry model of the flow cell (Figure S2). The fluid was assumed to be water under laminar-flow conditions. The flow rate was assumed to be 10 \u03bcL/min for the calculations. The model analyte was catechol (10\u03bcmol/L), which has frequently been used to evaluate microfluidic electrochemical detectors [33]. The diffusion coefficient was set to 2.2\u00d710 6 m2/ s, as experimentally determined by cyclic voltammetry. We designed three types of ring-array electrode (EQ, INC, and DEC), as shown in Figure 3. Detailed information regarding electrode geometry is depicted in Figure S3. These electrodes consisted of concentric ring electrodes with equal or different widths and constant spacings of 75 \u03bcm. The widths of the ring electrodes were a constant 50 \u03bcm for EQ, increased from 50 \u03bcm to 100 \u03bcm in increments of 10 \u03bcm from the inter to the outer ring for INC, and decreased from 100 \u03bcm to 50 \u03bcm in increments of 10 \u03bcm for DEC. A 1-mm-diameter conventional disc electrode (DISC) was employed for comparison" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002873_icpes47639.2019.9105570-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002873_icpes47639.2019.9105570-Figure5-1.png", + "caption": "Fig. 5: Rotor models with different numbers of pole pairs p.", + "texts": [ + " The stator is modelled with QS = 6 slots and wound with a double-layer concentrated winding (q = 0.5); the wound winding pattern is shown in the appendix in Fig. A-1. For a fair comparison the same stator is combined with rotors with different number of pole pairs. Thus depending on the rotor, the same geometry does or does not contain subharmonics. Changing the rotor is feasible for this winding scheme as the winding factor is the same for each harmonic (\u03b6\u03bd = 0.866). The rotor models are shown in Fig. 5. When changing the number of pole pairs of the rotor, the rotor speed has to be adjusted to the electrical stator speed with respect to the harmonic which is used to produce torque. \u2126S,\u03bd = \u03c9 \u03bd (10) \u2126R,\u03bd = \u03c9 \u03bd \u2212 \u03c9 p (11) In this section, the effects of the harmonics on the losses are presented. First, the effect of the single harmonic orders is considered. Secondly, the amplitude reduction of the MMF harmonics is shown. In the end, the impact of the subharmonics is explained. A. Influence of the harmonics 1) Fundamental and Harmonic Waves: In order to evaluate the impact of the fundamental wave and the single harmonics on the losses in the stator, rotor and PMs, a non-distributed integer-slot winding (q = 1) is considered", + " The rotor and PM losses are nearly on the same level, which can also be seen in Table II, where the values of the loss calulation of the fully- and fractionalpitched windings are listed. The sub-harmonics of the MMF distribution in the middle of the air gap have a substantial influence on the losses in the stator, rotor and PMs (Fig. 8). In order to better understand the sub-harmonic induced losses, an analysis of different models with different sub-harmonics is conducted. First, the sub-harmonics that occur by changing the number of pole pairs of the rotor model (Fig. 5) are considered. Secondly, the sub-harmonics of the single- and double-layer fractionalslot concentrated windings are analysed. 1) Changing Number of Pole Pairs: For the first study the stator model as shown in Fig. 4 with a number of slots QS = 6 is used. The corresponding winding pattern is shown Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 14,2020 at 15:36:02 UTC from IEEE Xplore. Restrictions apply. in Fig. A-1, and the rotor models in Fig. 5. The MMF spectra for all winding patterns are presented in Fig. 13. It can be seen that the MMF of the models with a pole number of p = 2 has no sub-harmonics. The winding scheme with p = 4 pole pairs has one and the model with p = 8 pole pairs has two sub-harmonics. These harmonics have a significant influence on the losses which can be seen in the loss diagram in Fig. 14. The sub-harmonics have an effect not only on the stator, but also on the rotor and PM losses. All of the losses increase dramatically with the number of poles which can be seen in detail in Table III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003461_embc44109.2020.9175945-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003461_embc44109.2020.9175945-Figure4-1.png", + "caption": "Fig. 4. Inside of the operating interface", + "texts": [ + " We used a basic structure of the three fingers developed by our research group in the past [12]. Each finger is connected to the center connector part with one joint. Three fingers are simultaneously opened by pushing the connector with the plunger with a stroke of 10 mm. The polyacetal guide pin of the finger moves in the guide track, which controls the trajectory of the finger. Conversely, three fingers are closed by returning the connector to the initial position by two extension springs (spring constant 0.49 N/mm) and one compression spring (spring constant 4.1 N/mm). Fig. 4 shows the inside of the operating interface. When the lever head is pushed by a user, the lever rotates around a rotation axis, and then the plunger of the small-diameter syringe is pushed. Three fingers are opened by pushing the lever of the operating interface, as shown in Fig. 5. The operating interface is mounted to the user\u2019s upper arm on the affected side with a band, and the user operates the lever by pinching it between the upper arm and the side, as shown in Fig. 6. The operating interface can be worn without a harness, unlike the body-powered prosthetic hand" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003835_biorob49111.2020.9224433-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003835_biorob49111.2020.9224433-Figure2-1.png", + "caption": "Fig. 2. The definition of knee angle \u03b8k , thigh angle \u03b8th, and shank angle \u03b8sh.", + "texts": [ + " Basically, the task of the motion planner is to estimate a desired motion for the missing joins/limbs in line with the motions of the remaining joins/limbs. In this study, data of thigh angular velocity \u03b8\u0307th and thigh angle \u03b8th were collected for walking 0.6, 0.9, 1.2, 1.4 and 1.6 m s (healthy male subject, measurement was done through IMU from Xsens MTw Awinda, 100 Hz). The knee joint angles \u03b8k were obtained by measuring the shank angles and thigh angles and then converting them to the knee angles (\u03b8k = 180\u2212(\u03b8sh\u2212\u03b8th)). The definitions of these angles are schematically shown in Fig. 2. It should be noted that the motion planner did not require shank data for estimation. The shank data was only required to calculate the corresponding knee angles during walking experiments. These obtained (calculated) knee angles are later compared with the estimated (by the motion planner) ones to evaluate the estimation quality and the performance of the motion planner. As observed in Fig. 1, the thigh angular velocities \u03b8\u0307th and thigh angles \u03b8th are the inputs to the motion planner. The knee angles \u03b8k are the outputs of the motion planner" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003730_tmag.2020.3027291-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003730_tmag.2020.3027291-Figure4-1.png", + "caption": "Fig. 4. End magnetic field model of PMTM", + "texts": [ + " It can be seen from the above analysis that the source of the end-effect force for PMTM is mainly caused by the sudden change of the magnetic field at the end of the central worm stator. Therefore, it is necessary to analyze the end magnetic field of PMTM and the change of the magnetic induction intensity. To some extent, the positional relationship between the planet gear and the central worm stator can be regarded as the bending of permanent magnet linear motor. Before applying the Schwarz-Christoffel transformation (SCT), the following assumptions are made. The end magnetic field model of PMTM is shown in Fig. 4. (1) The length of the air gap is infinite. (2) The magnetic potential on the surface of permanent magnet teeth is \u03c60, and the magnetic potential at the center of the air gap is 0. The transformation function from Z plane to \u03c9 plane can be expressed as [11]. 1 1 [2 1 ln ] 2 1 1 Z (2) d 1 d 2 z (3) where \u03b4 is the air gap length between the planet gear and central worm stator. The magnetic induction intensity at the end of the central worm stator can be expressed as 0 0 0 2 = 1 B H (4) where \u03bc0 is the vacuum permeability" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001658_icems.2019.8921632-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001658_icems.2019.8921632-Figure8-1.png", + "caption": "Fig. 8. Magnetic flux density visualization of SM-MG: (a) HMM, (b) FEA.", + "texts": [ + " The inner air-gap and outer air-gap magnetic flux density distributions of SM-MG and ST-MG solved by FEA and HMM are shown in Fig. 4 \u2013 Fig. 7. For the SM-MG, the magnetic flux density obtained by HMM and FEA almost coincide. Notably, a little large difference can be observed for the inner air-gap magnetic flux density of ST-MG, as shown in Fig.7. That is because an equivalent method is adopted to transfer a rectangle into two sectors. Besides, a magnetic field distribution for all parts in SM-MG and ST-MG calculated via HMM and FEA are shown in Fig. 8 and Fig.9. The magnetic field distribution trends are the same for HMM results and FEA results. The strips in HMM results are caused by the assumption that the permeabilities in radial direction is a constant, which further causes error in the magnetic field distribution prediction. In addition, it can be seen that the outer air-gap magnetic flux density distribution of SM-MG is almost the same with that of ST-MG. This means that the shapes of PMs on the inner rotor have little influence on the magnetic field distribution after the modulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.16-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.16-1.png", + "caption": "Figure 1.16. Floating wheel. For a color version of this figure, see www.iste.co.uk/jaulin/robotics.zip", + "texts": [ + " In the case where the articulations are rotational joints (as is the case of the Staubli robot where the joins can turn), the parametrization represented by the figure might prove to be practical since it makes it easier to draw the robot. This transformation is the composition of four elementary transformations: (i) a translation of length r following z; (ii) a translation of length d following x; (iii) a rotation of \" around y and (iv) a rotation of ! (the variable activated around z). Using the figure for drawing the arms and the photo for the robot, perform a realistic simulation of the robot\u2019s movement. EXERCISE 1.12.\u2013 Floating wheel Consider a wheel floating and spinning in the space as shown on Figure 1.16. We assume that the wheel is solid, similar to a homogeneous disk of mass m = 1 and radius , = 1. Its inertia matrix is given by: I = $ 6% m'2 2 0 0 0 m'2 4 0 0 0 m'2 4 & 8' 1) Give the state equations of the wheel. The state variables we choose are (i) the coordinates p = (x, y, z) of the center of the wheel, (ii) the orientation (&, !,') of the wheel, (iii) the speed vr of the center of the wheel in the wheel frame, and (iv) the rotation vector !r expressed in the wheel frame. We thus have a system of order 12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001185_10402004.2019.1664685-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001185_10402004.2019.1664685-Figure2-1.png", + "caption": "Figure 2. Degrees of freedom of the FPs/SPs.", + "texts": [ + " 1), the friction pair and the spline are in mesh, the friction plates (FPs) are connected to the input shaft by the external splines with input torque from the engine, rotating around the Zaxis with the speed of w (rad/s), and the steel plates (SPs) are connected to the output shaft by internal splines. In this article, the dynamic behaviors of the friction pairs are studied when the wet clutch is in a braking condition, so the rotation speed of the SP is zero. Due to the clearance between the FP and SP and the assembly error gap between the friction pair and the spline, under the fluid action, the FPs and SPs will translate along the Z-axis with linear displacement z and oscillate around the X- and Y-axes with angular displacements a and b, respectively (see Fig. 2). The shafts will hinder the angular deflections of the FPs and SPs. Previous studies were mainly focused on the drag torque caused by viscous shear of lubricating oil (Hu, et al. (1); Yuan, et al. (2); Wu, et al. (3); Pahlovy, et al. (4); Zhang, et al. (5)), but little was known about the drag torque phenomenon caused by the impact of friction pairs at high circumferential velocities. Mantwill found (6) Impact in a wet clutch will result in a sharp increase in drag torque. Therefore, the motion law of the friction pairs needs to be revealed to explain the impact process and then reduce the drag torque caused by impact in a wet clutch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001296_aim.2019.8868485-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001296_aim.2019.8868485-Figure5-1.png", + "caption": "Figure 5. Continuum robot structure design", + "texts": [ + " In this paper, the sensitive element is a hemispherical shell structure made of pressure-sensitive rubber, and several electrodes are arranged at equal intervals on its edge. These electrodes are used to exert excitation on the sensitive material, and measure the edge potential of the sensitive material. The electrodes are connected to the excitation circuit and data acquisition circuit through a multiplexer, and the detected edge potential is transmitted to the computer for calculation after data acquisition, so as to reconstruct the resistivity distribution diagram inside the sensitive element and obtain the position of the contact point. As shown in Fig.5, the structure design of continuum robot should meet the requirements of flexibility. In order to verify the performance of the continuum robot prototype, the continuum robot designed in this paper is a three-segment continuum structure. Each segment is controlled by four driving ropes distributed on the circumference. The two opposite driving ropes are in a group and driven by one motor. The body of the continuum robot is composed of a series of connecting disc. In order to ensure the good stiffness characteristics of the continuum robot, two adjacent connecting discs are connected with nickel-titanium alloy wires" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001369_012021-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001369_012021-Figure1-1.png", + "caption": "Figure 1. An example of thermal expansion leading to contact opening. t0 is the shape of the body at the initial temperature, t1 is the deformed warmed body with the open contact: a) the contact opening upon buckling due to thermal expansion; b) opening due to the deformation of the samples in the thermal contact conductance measurement test.", + "texts": [ + " This expansion in combination with the constraints of connections in the assembled structure leads to increased stresses and additional deformation. Thus, both tightening and gapping of contacting parts are possible. There are possible contact options, when the thermal contact conductance does not reach its steady-state value for a long time or even never reaches one if it has an oscillatory character. This is possible if thermal expansion makes the contact open. For example, the controlled process of mechanical contact opening can be seen in widespread constructions of thermostats and thermoregulators. Figure 1 shows possible simplified construction of systems, where the transient process may fluctuate. ICIPCE 2019 IOP Conf. Series: Materials Science and Engineering 630 (2019) 012021 IOP Publishing doi:10.1088/1757-899X/630/1/012021 In real structures of complex shape there will be an intricate and often unpredictable surface warp, also influenced by dependence on heat fluxes and compliance resulting from the asperity deformation. Such a warp should be taken into account by solving transient problems with finite element models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002299_rteict42901.2018.9012428-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002299_rteict42901.2018.9012428-Figure7-1.png", + "caption": "Fig. 7. Magnetic Parameters of 25% Short LSPMSM. (a) Radial Flux Density. (b) 2-D Flux Density Distribution. (c) Spatial FFT of Radial Flux Density distribution", + "texts": [ + " In this paper, stator turn-to-turn short-circuit faults of 25% and 50% magnitude are introduced in phase A of the machine by cutting out 104 turns (for 25% short) and 208 turns (for 50% short) out of a total number of 416 turns per phase. A. Parameters of 25% short LSPMSM: The faulty machine is simulated from Simplorer under the same supply and load conditions. The shorted phase current shows an increase in magnitude from 3.2A peak to 5A peak and decrease in induced voltage as depicted in Fig. 5. Due to the incorporation of fault, the speed and torque ripple increases in the machine which can be seen in Fig. 6(a) and Fig. 6(b). From Fig. 7(a), the un-symmetry in the radial flux density of the faulty machine can be noticed. The saturation of flux density in some parts of the machine due to short-circuit is encircled in Fig. 7(b). The spatial FFT of airgap flux density is shown in Fig. 7(c), which shows the introduction of noise content in the FFT spectrum due to asymmetry created by short-circuit fault compared to healthy condition. B. Parameters of 50% short LSPMSM: The 50% shorted phase current shows a further increase in magnitude of nearly 9A peak and decrease in induced voltage as depicted in Fig. 8. After the incorporation of 50% inter-turn fault, the speed and torque ripples increases significantly compared to 25% short-circuit condition as depicted in Fig. 9(a) and Fig", + " The results of co-simulation of healthy and faulty LSPMSM indicate that as the intensity of stator turn-to-turn short-circuit fault increases, the magnitude of stator current in the corresponding shorted phase increases proportionately. Also, oscillations can be noted in the moving speed of faulty machine leading to increased torque ripple. The radial flux density distribution becomes more asymmetrical and distorted as the fault condition degenerates. The flux distortion can be observed from the 2-D geometrical view of the faulty machines in the regions encircled black in Fig. 7(b) and Fig. 10(b). Also the spatial FFT of radial flux density shows that the flux density of the faulty machine is more polluted with sub-harmonics when compared to that of the healthy machine. Parameters which show measurable changes when there is a stator turn-to-turn SC fault are Stator Current, Machine Induced Voltage and Radial air-gap Flux Distribution. Among these, stator current is the simplest quantity which is easily accessible and directly available for analysis. Hence MCSA can be a suitable fault detection method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003322_tie.2020.3013520-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003322_tie.2020.3013520-Figure4-1.png", + "caption": "Fig. 4. (a) Phasor diagram of the IG system feeding CVVF AC loads with load voltage as reference phasor; (b) Simplified Phasor diagram with load current as reference phasor; (c) Phasor diagram of the IG system supplying both DC-link loads and CVVF AC loads. Ireal and Ireactive represents the real and reactive components of stator current, I1.", + "texts": [ + " Reactive power is supplied by the VSI which acts as a series compensator, while the active power needed for system operation is provided by the prime mover of the IG. Of the active power generated by machine, a major share will be delivered to the AC loads, and a part of the power is consumed by the inverter losses. During steady state, the real-power output (at DC-link) of the inverter should be zero, a capacitor fed VSI, with regulated DC-link voltage can ensure the zero DC-power condition. The phasor diagram for the proposed open-end winding IG in mode1 operation is shown in Fig. 4a. Vs is the converter voltage which is the sum of rotor induced voltage and voltage drop of stator resistance (including load resistance) and the leakage reactance. I1 is the stator current, and eb represents the backemf of the machine. Since the active power consumed by the inverter is nearly zero, I1 lags inverter AC output voltage by almost 90\u25e6. In mode-2 both DC-link loads and CVVF AC loads are connected to the open-end winding IG system and the active power generated by the induction machine is delivered to both AC and DC loads. Since the reactive power required by the machine is provided by the VSI, the instantaneous active power through the inverter will be supplied to the DC loads also. In this scheme, DC load is assumed to operate at nearly constant DC-link voltage as in the case of battery chargers. Thus maintaining a constant Vdc allows the DC load to draw the required power from the DC-link of the inverter. The phasor diagram for the proposed open-end winding IG in mode-2 with both DC and AC loads is shown in Fig. 4b where eb is the back emf, Ireal and Ireactive are the active and reactive part of stator current [29]. The complete system model is shown in Fig. 5. The dynamic model for the proposed system including open-end winding induction machine is required for simulation, transient analysis and controller design. The conventional induction machine model, which is a voltage-input current-output model, cannot be used for modeling the proposed system since the machine is operated as an open-end winding induction machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001754_icasert.2019.8934515-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001754_icasert.2019.8934515-Figure10-1.png", + "caption": "Fig. 10: Servo Mechanism.", + "texts": [], + "surrounding_texts": [ + "A portion of the \u2018DOORMOR Controller\u2019 app has the necessary tools to control the arm movement. There are three sliders for controlling the three servos [14]. The sliders have values from 0\u00b0 to 180\u00b0. These sliders need to be moved in a systematic way in order to obtain desired arm movements. The status area of the control app shows in which degree the servo [14] is at, whenever the concerned slider is highlighted. On the programming side, these sliders have different predefined underlying values at different points, e.g. the first slider has a value range of 1000-1180, the second slider has a value range of 2000-2180 and the third one has a value range of 3000-3180. These values are sent as strings which are further converted to integers and stored inside arrays in the microcontroller. The first digit represents the servo number and the rest three digits represent the degree at which the servos need to rotate and stay. After receiving signals from the control device, the microcontroller generates appropriate Pulse Width Modulation (PWM) signals for driving the servos [14]. PWM signal generation is handled by the Servo library of the Arduino language. C. PIR Sensor Status Another portion of the \u2018DOORMOR Controller\u2019 app displays the status of the PIR sensor [20]. The PIR sensor [20] can detect Passive Infrared Rays (PIR) emitted from living beings, such as, alive humans and other animals [28]. A dead body doesn\u2019t emit such rays. Inside the PIR sensor, all the thermal energy is at first focused on to the Pyroelectric sensor through the focusing lens in front of the sensor. If PIR is available, the voltage across the sensor changes. This change in voltage is too delicate to be detected by the microcontroller, so we need an amplifier circuit which will amplify the change. The output of the amplifier is connected to a voltage comparator, which compares voltages from the amplifier with a reference voltage and generates HIGH/LOW logic [28]. So, the module sends a HIGH logic to the microcontroller whenever it detects a living being within 1 meter. The microcontroller then sends a signal to the control device through the Pi and the Wi-Fi network. This signal is translated in the app. The app shows \u201cHuman Detected\u201d in the sensor status area, whenever a living body is detected. If there\u2019s no living body, the app shows \u201cNo Human Detected\u201d. The mechanism of the PIR sensor [20] is shown in Fig. 13." + ] + }, + { + "image_filename": "designv11_80_0001687_icems.2019.8922312-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001687_icems.2019.8922312-Figure1-1.png", + "caption": "Fig. 1. Machine structures. (a) PMVM. (b) DPMEM.", + "texts": [ + " The purpose of this paper is to quantitatively compare two typical FMPMs, namely a regular surface-mounted PMVM [3] and a DPMEM evolved from the regular surfacemounted PMVM for fair comparison. In Section II, structures of the PMVM and DPMEM are described. In Section III, working principles of the two machines are elaborated. In Section IV, performance of the two machines is quantitatively compared by using 2D finite element method (FEM). Finally, conclusions are drawn in Section V. II. STRUCTURE DESCRIPTION Fig. 1(a) shows the structure of a regular surfacemounted PMVM. Rotor PMs with alternate poles are surface-mounted on the rotor. The number of rotor PMs is 2Z2, and the pole pair number (PPN) of the rotor PMs is Z2. The stator has a tooth-pole structure [14], where stator teeth and wide slot-openings are distributed along the circumference of air-gap in order to utilize the fieldmodulated effect (FME). The numbers of stator tooth and stator slot are both Z1. The 3-phase armature windings are deployed in the stator slots. The slot wedges with epoxy resin material are employed to prevent coils from dropping. Fig. 1(b) shows a DPMEM evolved from Fig. 1(a) for a fair comparison in Section IV. Referring to Fig. 1(a), half rotor PMs with the same magnetization direction (MD) are retained and the other half rotor PMs are replaced by rotor teeth. Meanwhile, the material of slot wedges is replaced by PM material. Thus, one can obtain the DPMEM in Fig. 1(b). As can be seen from Fig. 1(b), both the stator and rotor have the tooth-pole structures, which means that both of them have the FME. The stator PMs and rotor PMs are surface-This work was supported by the Science and Technology Innovation Committee of Shenzhen under Projects ZDSYS201604291912175. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE inserted into slots, all PMs have the same MD, and each PM and its adjacent iron tooth form a pair of magnet poles. Therefore, the numbers of stator slots and stator teeth, PPNs of the rotor PMs and stator PMs are Z1, Z1, Z2, and Z1, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002138_cac48633.2019.8996404-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002138_cac48633.2019.8996404-Figure1-1.png", + "caption": "Fig. 1: Model of the quadruped robot", + "texts": [ + " This paper is organized as follows: Section II describes the physical model of the quadruped robot and necessary parameters. Section III demonstrates the CPG approach to generate joint desired trajectories. Section IV addresses a fuzzy adaptive PID control method according to desired trajectories. Section V presents relevant simulation results. At last, the conclusions and future works are mentioned in Section VI. In this section, the model of the quadruped robot used in this paper is established and shown in Fig. 1, related physical parameters and their meanings are demonstrated in TABLE I. Actually, feet of the quadruped robot are only aimed at mimicking animals and not taken into consideration in Section III. In order to fit in the simplified model in Section III, mass of femur m and length of femur l are introduced in m = mt = ms +mf (1) l = lt = ls + lf (2) For the quadruped robot, one leg has three joints\u2014hip, knee and ankle joint. Here ankle joints of all legs are also not considered. The configuration of joints is shown in TABLE II. This paper only pays attention to angles of joint hi and ki, where i is the serial number of legs and 1, 2, 3 and 4 represent left-front, right-front, right-rear and left-rear leg respectively. Based on the model built in Fig. 1, there are two problems solved in the rest of the paper: 1) Generate desired trajectories of hip and knee joints in the quadruped robot (solved in Section III). 2) Realize tracking control given desired trajectories (solved in Section IV). In this section, CPG model is used in producing joint desired trajectories of quadruped robots. In neurobiology, CPGs are neural networks which can generate coordinated patterns of rhythmic activity without any rhythmic inputs from sensory feedback or from higher control centers [6]", + " Here we adopt triangle membership function for all vague language values which is shown in Fig. 4. Fuzzy rules [11] are specified in TABLE IV, TABLE V and TABLE VI. According to (24), the output variables that exert on PID controller are supposed to be accurate quantities. The process is called defuzzification. Here the centroid method is adopted [12]. We use MATLAB/SIMULINK to generate desired trajectories of joint angles as the set values for joint motor controllers in our quadruped robot, which is modeled in ROS and as shown in Fig. 1. The relevant parameters are displayed in TABLE VII and TABLE VIII. The desired trajectories of four GPs can be produced according to TABLE VIII. Note that (GP, v, T, h) determines (\u03c9sw, Ah, Ak, \u03bc). Here the walk and trot gait are given in Fig. 5. Since it\u2019s requisite for the CPG network to stabilize in a few periods, part of desired trajectories are selected which are stable from t = 5s to t = 8s. The simulation results of walk gait using conventional PID method and fuzzy adaptive PID method are compared in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000202_sustech.2018.8671386-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000202_sustech.2018.8671386-Figure8-1.png", + "caption": "Fig 8. Rigid frame for inhouse multi-material 3D printer", + "texts": [ + " These experiments provide the foundation of our evolving in-house 3D printer prototype which will incorporate two printing heads (one for aluminum and the other for silicone plastic). A milling head will be added to the 3D printer prototype to hybridise additive manufacturing with subtractive manufacturing for superior finishing and tolerances. However, we shall be employing several approaches to improve the tolerances of FDM to minimise the amount of milling required: (i) reinforcing the stiffness of the structure including platform rails and extrusion head support \u2013 this has already been implemented (Fig 8); (ii) damping of vibrations using antivibration mounts; (iii) optimising flow viscosity during extrusion. It has been recommended that tolerance factors i are benchmarked against ISO 286-1:1988 grades [25]. For example, a 1 mm sized feature requires a tolerance factor i=0.542 \u03bcm with IT5, IT7 and IT12 being defined as 7i, 16i and 160i respectively, i.e. 3.8 \u03bcm, 8.6 \u03bcm and 0.086 mm respectively. The prospect of manufacturing active electronics and computing devices from lunar material appears remote - the cost of the GlobalFoundries Facility near Albany New York was $12B and requires highly stringent processes and materials that would be difficult to implement on the Moon" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000505_14484846.2019.1626529-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000505_14484846.2019.1626529-Figure12-1.png", + "caption": "Figure 12. Comparison of vortex distributions in the films.", + "texts": [ + "It can be obtained that the energy consumption efficiency of the orifice aerostatic bearing is higher than that of the porous aerostatic bearing, but the difference between them diminishes with the increase of gas-film thickness. Especially, when the gas film thickness is 14 \u03bcm, the difference between them is less than 10% of the energy efficiency of the porous aerostatic bearing. (2) Comparison of dynamic performances of bearings In order to characterize the dynamic performances of bearings, the vortex distributions in the film are exported from the CFD software. It can be concluded from Figure 12 that the vortex in the orifice bearing is mainly concentrated around the orifice and the outlet of the gas chamber, while the vortex in the porous bearing is mainly distributed at the outer edge of the porous restrictor. From the perspective of bearing stability, the porous aerostatic bearing presents superior dynamic performance. In conclusion, in terms of the single-restrictor aerostatic bearing, the porous restrictor is preferred under the same operating conditions. 4.1. Optimization of the number of orifice restrictors In practice, the aerostatic bearing is generally configured with multiple restrictors to support the load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002315_012062-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002315_012062-Figure5-1.png", + "caption": "Figure 5. Strain gauge positions for gear load distribution evaluation [11].", + "texts": [ + " Any generic load can be applied and the obtained measurement results are highly reproducible. Figure 4 shows the research nacelle mounted on the system test bench at the CWD. Amongst roughly 300 sensors in this turbine, the measurement equipment included rotating strain gauges on the sun and stationary strain gauges on the ring gear at four positions over the circumference [10]. By this, the load distribution in both planetary gear\u2019s contacts can be evaluated over the gear width and over the ring gear\u2019s circumference. In figure 5, the strain gauge positions on the sun and ring gear are shown. Also, distance sensors were installed in the planet carrier, facing the planetary gear and its bearing outer rings from both sides [10]. In figure 6, the positions of the distance sensors used for the presented creep evaulation are shown. The described setup allows the axial positions of the planetary gear and its bearings\u2019 outer rings to be measured directly. By calculating the difference of these signals, axial creep can be detected", + " As ring creep is a very slow and most importantly cumulative process and literature shows no influence of dynamic loads on ring creep, dynamic loads were not investigated. In addition, the presented detection and evaluation method would not be able to process tangential creep phenomenon on very small time scales to see the influence of dynamic loads. This chapter presents the results of the chosen load scenarios with respect to gear load distribution and creeping behavior. For every test, gear load distribution was evaluated at four positions on the ring gear (see figure 5) and the gear load distribution of the sun gear was reduced to these areas for easier comparison. In figure 9 the gear load distribution for 40% and 100% nominal rotor torque is shown. For better overview, only the \u201c180\u00b0-(low) position\u201c as defined in figure 5 is shown. The blue crosses represent the sun gear measurement, the red crosses the ring gear measurement. One of the sun gear strain gauges had to be interpolated due to technical issues during data acquisition. The percentage values of each color are the sum of the three strain gauge measurements of this side (rotor or generator) divided by the sum of all six measurements of this component. Therefore, this value \u2013 in a simplified way \u2013 represents the load carrying share of this half of the tooth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001886_summa48161.2019.8947534-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001886_summa48161.2019.8947534-Figure2-1.png", + "caption": "Fig. 2. Manipulator scheme on coordinate plane.", + "texts": [ + " In order for this situation to exist, the following conditions must be satisfied, obtained from relations (5) - (11): -g m1 k1-a2 cos v1 0+m2(k1cosv1 0+k2sin v1 0++v2 0 + + L1 d1sin w1 0-v1 0 -a1cosv1 0 d1 L1sin w1 0-v1 0 -a1cosw1 k2(1)i10=0 (12) -gm2k2sin v1 0+v2 0 + L2 d2sin w2 0-v2 0 -a2c osv2 0 d2 L2sin w2 0-v2 0 -a2cosw2 0 k 2(2) i20=0 (13) R1i10=e1 0, R2i20= e2 0, (14) x0=k1cosv1 0+k2sin v1 0+v2 0 , (15) y0=k1sinv1 0-k2cos v1 0+v2 0 (16) L1 2+d1 2+a1 2-2L1d1cos w1 0-v1 0 +2L1a1sinv1 0-2a1d1sinw1 0=l1 2 (17) L2 2+d2 2+a2 2-2L2d2cos w2 0-v2 0 +2L2a2sinv2 0-2a2d2sinw2 0=l2 2,(18) From conditions (12) - (18) under the additional condition of choosing a larger value for v1 0 (as shown in Fig. 2), all constant values v1 0,v2 0,w1 0,w2 0,i10,i2 ,0,e1 0,e2 0, corresponding to the given values x0,y0can be uniquely determined. Introduce perturbations in the neighborhood of a given point x0,y0,v 1 0,v2 0,w1 0,w2 0,i10,i2 ,0,e1 0,e2 0 v1=v1 0+x1, v1=x2, v2=v2 0+x3,v2=x4, w2=w2 0+y2, e1=e1 0+u1, e2=e2 0+u2. Here u1,u2 - the additional voltage on the anchor windings of the drive motors will be taken for control actions. We pose the problem of holding the grip of the manipulator at a given point as the stabilization problem x1=x2=x3=x4=x5= x6=y1=y2=0 (19) III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure3.11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure3.11-1.png", + "caption": "Figure 3.11 The scheme of a combined system of the rigid-beam type and flexible-arch with the ride on the top.", + "texts": [ + " We made an attempt to determine the free (natural) oscillations of combined systems, taking into account deformations, not only of the lowest frequencies but also of a range of frequencies for a combined system, i.e. a flexible arch and rigid beam, containing the spectrum of frequencies [263, 280]. In these systems the interest lies in discerning in which cases the lowest frequency will be arch \u2013 reverse-symmetrical; and in which cases the lowest frequency will be beam \u2013 symmetrical. The obtained theoretical results were tested on models and structures (Figure 3.11). 3.8 The Free Oscillations of System \u201cFlexible Arch-Rigid Beam\u201d 183 Let us examine the free oscillations of this combined system [280]. The integro-differential equation of the oscillations takes the form: \ud835\udf154\ud835\udf02 dx4 + H EbIb \u22c5 \ud835\udf152\ud835\udf02 dx2 \u2212 \u0394H EbIb \u22c5 \ud835\udf152y dx2 + m EbIb \u22c5 \ud835\udf152\ud835\udf02 dt2 = 0 (3.8.1) This equation is obtained with the following prerequisites (preconditions): the axis of the arch is outlined on the square parabola; the posts are not under tensile force (inextensible, nonextensible) and they are not compressed; rigidity of the beam is constant on the entire span; and the mass of the bridge span is distributed along the length of the rigid beam" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003696_0954409720962245-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003696_0954409720962245-Figure5-1.png", + "caption": "Figure 5. Magnitude of OOR in the polar coordinate system. (a) Left wheels and (b) Right wheels.", + "texts": [ + "24 h\u00f0t\u00de\u00bcmod xt 360 (6) where x is the nominal angular velocity of the wheel and t is the travel time. The wheel radius at the wheel-rail contact point at time t during the test, R(t), can be expressed as R\u00f0t\u00de \u00bc R\u00feDr\u00f0t\u00de (7) where R is the nominal rolling radius of the wheel. The OOR data was acquired by measuring the irregularity in the nominal rolling circle along the wheel circumference. For the study to be more realistic, the eight wheels of a selected carriage were analyzed for their OOR. The measured OOR of these wheels is represented in the polar coordinate system, as shown in Figure 5. Figure 5(a) and (b) illustrate the magnitude of OOR for the four left wheels (numbered 1, 3, 5, and 7) and the four right wheels (numbered 2, 4, 6, and 8), respectively. It is clear from the figure that the radial runout of each wheel is different, which suggests obvious asymmetry of OOR. Therefore, it is unwise to define the same distribution of OOR in the wheels as being symmetric. Vehicle dynamics model featuring asymmetric OOR coupled with wear Dynamics model of vehicle system A vehicle dynamics model was created based on the parameters of a subway train in China", + " In this study, the radial deviation approach was adopted to set OOR for the vehicle dynamics model. In this approach, INPUT FUNCTION provided in SIMPACK was used to express OOR in polar coordinates as (x, f (x)), where x is the polar angle of a point on the wheel circumference, x 2 \u00bd0; 2p , and f (x) represents the wheel diameter (m). The spacing between data points along the wheel circumference should be as small as possible to ensure accuracy. Then the OOR data from the eight measured wheels shown in Figure 5 were applied to the eight wheels for modeling of asymmetric OOR.30 Later, the wheel wear data measured at different operational stages were imported to the wheel tread database. In asymmetric wear setting, the profile of the eight wheels for each distance traveled shown in Figure 2 was used to the eight wheels. After that, the wear data was applied to corresponding wheels in the simulation model to complete the modeling of asymmetric OOR coupled with asymmetric tread wear. Influence of OOR coupled with tread wear on vehicle dynamics Influence of asymmetric OOR on vehicle dynamics Previous research usually provided identical OOR profiles to all wheels for simplicity. Given the asymmetric distribution of OOR in actual operation, the vehicle dynamics model presented in Figure 8 and the asymmetric OOR model constructed in the above section were used for analysis of the influence of asymmetric OOR on vehicle dynamics and comparison with symmetric OOR. For the symmetric OOR model, the average OOR of the eight measured wheels in Figure 5 was applied to all wheels in order to give them the same OOR profile. The guide wheelset was used as an example to illustrate the subway vehicle\u2019s dynamic response, as its dynamic response was the strongest among the four wheelsets. Vertical wheel/rail contact force. In actual operation, a separation between wheel and rail is bound to cause a very strong wheel\u2013rail impact. When the separated wheel and rail come into contact again, the impact process will generate not only a great force, but also significant shock acceleration", + " To collect acceleration data, an accelerometer was installed on the left side of the leading bogie of each wheelset (1m apart from the bogie\u2019s center), a proper site for measuring train stability as required by Railway vehicles - Specification for Evaluation the Dynamic Performance and Accreditation Test (GB5599-1985).32 The sample frequency was 1024Hz, and each test wheel had traveled 5 104 km in total. Condition I: symmetric tread wear coupled with symmetric OOR, with the eight wheels\u2019 average wear (Figure 2) and average OOR (Figure 5) for the distance traveled of 5 104 km applied to all the eight wheels. Condition II: asymmetric tread wear coupled with asymmetric OOR, i.e. the aforementioned type B coupled damage (Table 3), with the tread wear and OOR data for the distance traveled of 5 104 km adopted for all wheels. All else equal, acceleration measurement points were laid out in the same manner under Conditions I and II. Then the simulated vertical acceleration was output. Figure 18 compares the measured vertical acceleration with the simulation results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003834_biorob49111.2020.9224305-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003834_biorob49111.2020.9224305-Figure1-1.png", + "caption": "Figure. 1. The sensor system employed in the current study", + "texts": [ + " The smaller target is often challenging for the LIDAR to detect if it is poorly reflective, distant, and its aspect is away from the normal. Since the current application is targeting humans, these limitations might be exempted here. The LIDARLITE v3HP module (2x1x0.5 inch) has a 2- wire, I2C-compatible serial interface which supports 400 kHz Fast Mode for data transfer. In this study, the LIDARLITE module is connected to the I2C bus (as a slave device) of a low-cost ATmega2560 microcontroller (I2C master device, Figure 1). ATmega2560 was clocked by a 16 MHz external crystal oscillator. This 8-bit AVR RISC-based microcontroller features to have 8 KB SRAM, 256 KB ISP flash memory, and an average throughput of 16 MIPS at 16 MHz clocks. Considering the maximum frame rate (1KHz) of LIDAR, the microcontroller was programmed to read distance data from LIDAR at 1KHz, store in the memory, and transfer to the computer via its serial (UART) interface for real-time data monitoring. To communicate with a computer, the UART baud rate was set as 115200 bps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003616_j.matpr.2020.08.536-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003616_j.matpr.2020.08.536-Figure2-1.png", + "caption": "Fig. 2. Worm and Wheel Assembly.", + "texts": [], + "surrounding_texts": [ + "A rack and pinion mechanism along with worm and wheel gear for producing electricity. The type of shock absorber used is pneumatic shock absorbers (Figs. 2\u20135). The detailed methodology is provided in Fig. 6. DC generator Rack arrangement Worm and wheel gear Pneumatic shock absorber" + ] + }, + { + "image_filename": "designv11_80_0001934_978-981-15-1124-0_17-Figure19-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001934_978-981-15-1124-0_17-Figure19-1.png", + "caption": "Fig. 19 Mode Shape 5", + "texts": [], + "surrounding_texts": [ + "shown in images Fig. 15 and Fig. 24. We can notice that the fracture of the femur bone seems to occur either at the shaft or the neck region for different frequencies. Although an external load can excite any mode from the first mode to any higher mode, the lower nodes are comparatively easier to excite." + ] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure15-1.png", + "caption": "Figure 15. Shear stress from a plain bearing to a textured surface (N = 6000rpm, W = 2000N)", + "texts": [], + "surrounding_texts": [ + "These\ufeff displacements\ufeff substantially\ufeff modify\ufeff the\ufeff radial\ufeff clearance\ufeff during\ufeff operation.\ufeff The\ufeff maximum\ufeff displacement\ufeffis\ufeffnoted\ufeffin\ufeffthe\ufeffangular\ufeffpositions\ufeffbetween\ufeff150\u00b0\ufeffand\ufeff200\u00b0.\nShear Stress Distribution Shear\ufeffstress\ufeffdistribution\ufeffin\ufeffmedian\ufeffplan\ufefftextured\ufeffplain\ufeffbearing\ufeffand\ufeffuntextured\ufeffplain\ufeffbearing\ufefffor\ufeffdifferent\ufeff radial\ufeffload,\ufeffis\ufeffpresented\ufeffin\ufeffFigure\ufeff15.\ufeffMaximum\ufeffshear\ufeffstress\ufeffis\ufeffrecorded\ufeffin\ufefftwo\ufeffangular\ufeffposition\ufeffof\ufeff the\ufeffplain\ufeffbearing\ufeff50\u00b0\ufeffto\ufeff70\u00b0\ufeffalso\ufeffat\ufeff220\u00b0\ufeffto\ufeff260\u00b0.\ufeffA\ufefftextured\ufeffplain\ufeffbearing\ufeffunder\ufeffa\ufeffload\ufeffof\ufeff10\ufeffkN\ufeffis\ufeff subjected\ufeffto\ufeffsignificant\ufeffshear\ufeffstress\ufeffcompared\ufeffto\ufeffa\ufeffnon-textured\ufeffplain\ufeffbearing.\nA\ufefftextured\ufeffplain\ufeffbearing\ufeffunder\ufeffa\ufeffload\ufeffof\ufeff10kN\ufeffand\ufeffa\ufeffrotational\ufeffspeed\ufeffof\ufeff6000rpm\ufeffis\ufeffsubjected\ufeffto\ufeff significant\ufeffshear\ufeffstress\ufeffcompared\ufeffto\ufeffa\ufeffnon-textured\ufeffbearing\ufeff(Figure\ufeff16).\ufeffThese\ufeffstresses\ufeffreach\ufeff40.105\ufeff Pa\ufefffor\ufeffa\ufefftextured\ufeffbearing\ufeffand\ufeffreach\ufeffonly\ufeff9,5.105\ufeffPa\ufefffor\ufeffa\ufeffnon-textured\ufeffplain\ufeffbearing.\nRotational Velocity Effect Pressure The\ufeffeffect\ufeffof\ufeffthe\ufeffrotational\ufeffspeed\ufeffof\ufeffthe\ufeffshaft\ufeffon\ufeffthe\ufeffpressure\ufeffdistribution\ufeffin\ufeffthe\ufeffplain\ufeffbearing\ufeffmedian\ufeff plane\ufeffof\ufeffthe.\ufeffThe\ufeffrotational\ufeffvelocity\ufeffof\ufeffthe\ufeffshaft\ufeffis\ufeffvaried\ufefffrom\ufeff2000\ufeffrpm\ufeffto\ufeff9000\ufeffrpm;\ufefffor\ufeffa\ufeffsupply\ufeff pressure\ufeffof\ufeffPa\ufeff=\ufeff0.04\ufeffMPa,\ufeffsupply\ufefftemperature\ufeffof\ufeffTa=\ufeff40\ufeff\u00b0C\ufeffand\ufeffa\ufeffradial\ufeffload\ufeffof\ufeffW\ufeff=\ufeff10000N.\ufeffThe\ufeff", + "graph\ufeffclearly\ufeffshows\ufeffthat\ufeffthe\ufeffincrease\ufeffin\ufeffthe\ufeffspeed\ufeffof\ufeffrotation\ufeffof\ufeffthe\ufeffshaft\ufeffleads\ufeffincreased\ufeffhydrodynamic\ufeff pressure,\ufeffthe\ufeffsignificant\ufeffvalues\ufeffare\ufeffnoted\ufefffor\ufeffa\ufeffspeed\ufeffof\ufeff9000\ufeffrpm,\ufeff(Figure\ufeff17\ufeffand\ufeffFigure\ufeff18).\nDisplacement Distribution for Textured Plain Bearing The\ufeffmaximum\ufeffdisplacement\ufeffdue\ufeffto\ufeffthe\ufeffhydrodynamic\ufeffpressure\ufefffield,\ufeffis\ufeffnoted\ufefffor\ufeffan\ufeffangular\ufeffposition\ufeff of\ufeff150\ufeff\u00b0\ufeffto\ufeff200\ufeff\u00b0,\ufeffas\ufeffwell\ufeffas\ufeffit\ufeffis\ufeffinfluenced\ufeffby\ufeffthe\ufeffincrease\ufeffof\ufeffthe\ufeffrotational\ufeffvelocity\ufeff(Figure\ufeff19).\ufeff increasing\ufeffthe\ufeffspeed\ufeffof\ufeffrotation\ufeffof\ufeffthe\ufeffshaft\ufeffleads\ufeffto\ufeffthe\ufeffincrease\ufeffin\ufeffdisplacement.\nThe\ufeffmaximum\ufeffvalue\ufeffof\ufeffdisplacement\ufeffis\ufeffnoted\ufefffor\ufeffa\ufeffspeed\ufeffof\ufeff12000rpm\ufeffin\ufeffthe\ufeffcase\ufeffof\ufeffa\ufefftextured\ufeff plain\ufeffbearing\ufeffis\ufeffof\ufeffthe\ufefforder\ufeffof\ufeff0.39\u03bcm.\ufeffOn\ufeffthe\ufeffother\ufeffhand,\ufefffor\ufeffa\ufeffplain\ufeffbearing\ufeffwithout\ufefftexture,\ufeffthe\ufeff maximum\ufeffvalue\ufeffof\ufeffthe\ufeffdisplacement\ufeffis\ufeffof\ufeff0,32\u03bcm,\ufeffFigure\ufeff20.\ufeffThe\ufeffdisplacement\ufeffdistribution\ufeffallowing\ufeff to\ufeffangular\ufeffposition\ufefffor\ufeffdifferent\ufeffrotational\ufeffvelocity\ufeff(6000,\ufeff9000\ufeffand\ufeff12000rpm),\ufefffor\ufeffplain\ufeffbearing\ufeffat\ufeff textured\ufeffsurface\ufeffand\ufeffuntextured\ufeffplain\ufeffbearing\ufeffis\ufeffillustrated\ufeffin\ufeffthe\ufeffFigure\ufeff21.\ufeffthe\ufeffmaximum\ufeffdeformation\ufeff is\ufeffnoted\ufeffat\ufeffthe\ufefflevel\ufeffof\ufeffthe\ufefflower\ufeffgeneratrix,\ufeffas\ufeffwell\ufeffas\ufeffthe\ufeffdisplacement\ufeffvariation\ufeffof\ufeffbearing\ufeffwith\ufeffthe\ufeff rotational\ufeffvelocity\ufeffvariation\ufeffis\ufeffsignificant\ufefffor\ufeffa\ufeffplain\ufeffbearing\ufeffwith\ufeffa\ufefftextured\ufeffsurface\ufeffcompared\ufeffto\ufeffthat\ufeff obtained\ufefffor\ufeffa\ufeffnon-textured\ufeffplain\ufeffbearing.\nShear Stress The\ufeffshear\ufeffstress\ufeffevolution\ufeffaccording\ufeffto\ufeffangular\ufeffposition\ufefffor\ufeffdifferent\ufeffrotational\ufeffvelocity\ufefffor\ufeffplain\ufeff bearing\ufeffat\ufefftextured\ufeffsurface\ufeffand\ufeffuntextured\ufeffplain\ufeffbearing,\ufefffor\ufeffthe\ufeffradial\ufeffload\ufeffis\ufeff6000N,\ufeffis\ufeffillustrated\ufeff in\ufeffthe\ufeffFigure\ufeff18.\ufeffThe\ufeffmaximum\ufeffstresses\ufeffshear\ufeffvalues\ufeffare\ufeffnoted\ufeffin\ufeffthe\ufeffboth\ufeffangular\ufeffpositions\ufeff60\u00b0\ufeffand\ufeff 230\u00b0\ufeff(Figure\ufeff22).\ufeffThe\ufeffincrease\ufeffin\ufeffthe\ufeffrotational\ufeffvelocity\ufeffof\ufeffthe\ufeffshaft\ufeffleads\ufeffto\ufeffan\ufeffincrease\ufeffin\ufeffthe\ufeffshear\ufeff stress,\ufeffFigure\ufeff23.\ufeffThis\ufeffincrease\ufeffreaches\ufeff23\ufeffpercent\ufefffor\ufeffa\ufefftextured\ufeffplain\ufeffbearing,\ufeffon\ufeffthe\ufeffother\ufeffhand\ufefffor\ufeff a\ufeffnon-textured\ufeffplain\ufeffbearing\ufeffreaches\ufeff23\ufeffpercent." + ] + }, + { + "image_filename": "designv11_80_0001046_acc.2019.8814729-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001046_acc.2019.8814729-Figure1-1.png", + "caption": "Fig. 1. Sensor Configuration for position estimation of EHA", + "texts": [ + ". INTRODUCTION Magnetic position estimation offers an excellent inexpensive and non-contacting method of obtaining piston position in many modern actuators, including pneumatic cylinders, hydraulic actuators and IC engines. In magnetic position estimation, a magnet is placed on the moving object, such as the moving piston shown in Figure 1. A sensor board containing one or more magnetic sensors is placed on the outside cylinder, again as shown in Fig. 1. Such magnetic sensors are inexpensive (as low as $1 each when purchased in large quantities). At the same time, they enable non-contact estimation of position of the piston. Traditional sensors such as potentiometers and LVDTs require the sensor to be connected co-axially to the moving piston. This requires significant installation effort, results in contacting motion and in shear loads on the sensor during operation, often resulting in sensor failure. Furthermore, potentiometers and LVDTs can be more expensive than the low-cost magnetic sensors considered in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003544_ccc50068.2020.9188426-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003544_ccc50068.2020.9188426-Figure1-1.png", + "caption": "Fig. 1: Body coordinate system and inertial coordinate system", + "texts": [ + " Section three introduces the PID gain adjustment scheme using the PPO algorithm. Section four compared with the classic PID controller to demonstrate the performance improvements of the PID controller trained with PPO and conclusions are drawn in Section five. A quadcopter is an aircraft with four motors using a propeller propulsion system. It has six degrees of freedom (DOF), three rotational and three translational. The quadrotor with a cross-shaped symmetrical structure is chosen to show the performance of the PPO scheme as depicted in Fig. 1. 6756 Authorized licensed use limited to: Auckland University of Technology. Downloaded on November 03,2020 at 06:31:51 UTC from IEEE Xplore. Restrictions apply. The dynamics of a quadrotor have been investigated by many researchers. According to Newton's second law, the dynamic model of a quadrotor with respect to the inertial coordinate system can be defined as 1 1 1 2 3 4 (cos sin cos sin sin ) / (cos sin cos sin cos ) / cos cos / ( ( ) ) / ( ( ) ) / ( ( ) ) / x y z x y z y z x z x y x y z x v y v z v v U m v U m v U mg m p q r p lU I I qr I q lU I I pr I r U I I pq I (1) Where [ , , ]Tx y z represents the position of the quadrotor in the inertial coordinate; [ , , ]T is the attitude angle in the body coordinate system; m is the mass of the quadrotor; [ , , ]Tp q r represents the angular velocity of the quadrotor; [ , , ]Tx y zI I I is the moment of inertia matrix; 1 2 3 4[ , , , ]TU U U U U is the control input which is determined by the rotation speeds of the motors 1 2 3 4[ , , , ]T , and 1 2 3 4, , ,U U U U represent the thrust control amount, roll torque control amount, pitch and yaw torque control amount" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002493_042027-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002493_042027-Figure3-1.png", + "caption": "Figure 3 Sectional view of the equivalent stress of the ball raceway", + "texts": [ + " During the point contact process, Hertz established the following normal stress mathematical model, and the normal stress distribution of the contact ellipse is shown in the Figure 2: Where Q is the normal load, a is the long radius of the stress ellipse, and b is the short radius of the stress ellipse,\u03c3 Is the normal stress of the contact ellipse.It is easy to get the maximum pressure center to appear at the geometric center. During the loading process, due to the friction between the gap and the contact surface, the contact angle between the ball and the raceway changes during the actual loading process. As shown in Figure 3, the maximum stress surface of the ball raceway is misaligned, indicating that the raceway and the ball have shifted from each other during the loading process. The contact deformation and clearance of the ball and the upper and lower raceways will affect the ball. Actual contact angle. SAMSE 2019 IOP Conf. Series: Materials Science and Engineering 768 (2020) 042027 IOP Publishing doi:10.1088/1757-899X/768/4/042027 From Figure 4, it can be seen that when the displacement load is 0.1mm, the maximum equivalent stress is less than the yield strength of the ball raceway material", + " The position of the effect force is on the sub-table, and the stress depth is constantly increasing; when the ratio is greater than 0.1925, the von-miss maximum equivalent stress position is on the sub-table. 4.Conclusions When the support undergoes plastic deformation, the equivalent friction coefficient of the contact surface decreases with increasing displacement load. When the equivalent friction coefficient decreases to 0.1925, the maximum stress of the support occurs on the secondary surface, that is, the contact surface The following positions. (1)In the plastic stage, it can be observed from Figure 3 that the maximum equivalent stress appears SAMSE 2019 IOP Conf. Series: Materials Science and Engineering 768 (2020) 042027 IOP Publishing doi:10.1088/1757-899X/768/4/042027 on the subsurface of the ball and raceway, so it can be determined that the fatigue failure of the ball occurs from the inside of the material, and the cracks propagate from the inside to the outside. Causes pitting on the ball surface. (2) Analysis of the effect of the equivalent friction coefficient on the distribution of subsurface stress is conducive to the reasonable setting of processes such as carburizing, locally strengthening the raceway, and prolonging the service life of the slewing bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002158_cac48633.2019.8997353-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002158_cac48633.2019.8997353-Figure1-1.png", + "caption": "Figure 1", + "texts": [ + " Based on the data collected by gyroscope and the stability theory of robot, the reference trajectory is planned online. The output of predictive control is the elongation of the virtual leg. By tracking the reference trajectory, better stability can be obtained. The reliability of the algorithm is verified by virtual simulation and a prototype of a tripod robot. II. THREE-LEGGED ROBOT PROTOTYPE STRUCTURE In general, the multi-legged robot is inspired by spiders [11]. However, the three-legged robot mentioned in this paper is different from the ordinary multi-legged robot. The design structure is shown in Figure 1. The unit of the number marked in the figure is Millimeter. The fuselage body is made of aluminum alloy material and it is a disc structure. Three metal parts are connected to the disc, and a servo motor is mounted at the joint of the device and the disc, so that the motor can drive the metal device to rotate freely around the connecting shaft. The legs are made of carbon fiber tubes, and a servo motor is installed between the thighs and the metal parts to form a hip joint. A servo motor is mounted between the thigh and the lower leg to form a knee joint" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002886_apj.2510-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002886_apj.2510-Figure1-1.png", + "caption": "Figure 1 shows the squeeze film geometry of the problem. Consider an axially symmetric, incompressible, steady, and radial flow of an electrically conducting lubricant fluid squeezed between two circular disks which are separated a distance d apart. A uniform external magnetic", + "texts": [ + " The boundaries of the plug core thickness are given by z= d1 r\u00f0 \u00de and z= d2 r\u00f0 \u00de as illustrated in Figure 2. Based on the assumptions underlying in the lubrication system for thin films, the basic momentum equations governing the squeeze film motion are expressed as 1 r \u2202 r v\u00f0 \u00de \u2202 r + \u2202 u \u2202 z =0, \u00f02\u00de \u03c1 v \u2202 v \u2202 r + u \u2202 v \u2202 z = \u2212 \u2202 p \u2202 r + \u2202 \u03c4 \u2202 z \u2212 \u03c3 B2 0 v, \u00f03\u00de \u2202 p \u2202 z =0, \u00f04\u00de where u represents the axial velocity, \u03c1 denotes the density, p indicates the pressure, \u03c3 is the electrical conductivity, and B0 is the external magnetic field. The boundary conditions are FIGURE 1 Geometry of the squeeze film bearing between circular disks v=0at z=0, d, \u00f05\u00de \u2202 v \u2202 z =0 at z= d1 r\u00f0 \u00de and z= d2 r\u00f0 \u00de, \u00f06\u00de u=0 at z= 0, \u00f07\u00de u= \u2212 v0 at z= d, \u00f08\u00de p= p0 at r= R, \u00f09\u00de where p0 denotes the atmospheric pressure. The continuity Equation 2 in the integral form may be obtained as \u00f0 d 0 vd z= r v0 2 : \u00f010\u00de As suggested by Hashimoto and Wada,6 the integral form of the momentum Equation 3 using the continuity Equation 2 and the boundary conditions 5 to 8, we get \u03c1 d \u2202 \u2202 r \u00f0 d 0 v\u00f0 \u00de2d z+ 1 r \u00f0 d 0 v\u00f0 \u00de2d z 2 64 3 75+ \u2202 p \u2202 r + \u03c3 B2 0 d \u00f0 d 0 vd z= \u2202 \u03c4 \u2202 z : \u00f011\u00de By introducing the following analogous pressure gradient E= \u03c1 d \u2202 \u2202 r \u00f0 d 0 v\u00f0 \u00de2d z+ 1 r \u00f0 d 0 v\u00f0 \u00de2d z 2 64 3 75+ \u2202 p \u2202 r + \u03c3 B2 0 d \u00f0 d 0 vd z, \u00f012\u00de the above Equation 11 becomes \u2202 \u03c4 \u2202 z = E: \u00f013\u00de On integrating Equation 13 with respect to z, we have \u03c4= E z+ c: \u00f014\u00de Substituting Equation 14 into Equation 1, integrating the resultant equation and using the boundary conditions 5 and 6, we may obtain the velocity profiles in the two flow regions as v= E 1\u2212\u03bb\u00f0 \u00de k 1 m m m+1 z\u2212 d1 m+1 m \u2212 \u2212 d1 m+1 m h i , \u00f015\u00de in 0\u2264 z\u2264 d1 r\u00f0 \u00de, v= E 1\u2212\u03bb\u00f0 \u00de k 1 m m m+1 z\u2212 d2 m+1 m \u2212 d\u2212 d2 m+1 m h i , \u00f016\u00de in d2 r\u00f0 \u00de\u2264 z\u2264 d: The plug core velocity ( vp ) in the domain d1 r\u00f0 \u00de\u2264 z\u2264 d2 r\u00f0 \u00de may be obtained as vp = E 1\u2212\u03bb\u00f0 \u00de k 1 m m m+1 \u2212 d1 m+1 m h i = E 1\u2212\u03bb\u00f0 \u00de k 1 m m m+1 d\u2212 d2 m+1 m h i : \u00f017\u00de Using Equation 17, the following relationship is obtained as d1 r\u00f0 \u00de= d\u2212 d2 r\u00f0 \u00de: \u00f018\u00de The relationship between the analogous pressure gradient ( E\u00de and the yield value of the fluid is established as follows: E= \u2212 2 \u03c4y \u03b2 r\u00f0 \u00de , \u00f019\u00de where \u03b2 r\u00f0 \u00de= d2 r\u00f0 \u00de\u2212 d1 r\u00f0 \u00de: \u00f020\u00de It is noticed that \u03b2 r\u00f0 \u00de denotes the plug core thickness" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001369_012021-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001369_012021-Figure2-1.png", + "caption": "Figure 2. Geometry of macromodel (cross section shown, dimensions in mm). Sg-p is the contact surface of gyro unit\u2013platform.", + "texts": [ + " In fact, thermal contact conductance varies at different points of the contact and to solve the problem we calculate the dependence of thermal contact conductance on contact pressure with our column micromodel of rough surfaces contact [7]. Then, this dependence is transferred to the finite element model of the gyro unit\u2013platform assembly. 2. Description of gyro unit\u2013platform model and formulation In the right-handed Cartesian reference system, let us consider the hypothetical construction of a gyro unit attached to a platform (see Figure 2). The model of the gyro unit is a solid cylinder with a flange, made of AISI 1020 steel. Inside the gyro unit we assume the presence of a thermostatic system maintaining on a constant level the temperature of some internal part of the device. At the same time, the temperature of the device body is not constant and is determined, among other things, by interaction with the platform. The model does not consider the internal parts and the thermostatic system as separate bodies, but it takes into account that this system generates heat with a power of 20 W, corresponding to a specific heat generation rate q = 94832 W/m3 for the given volume of the gyro unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000530_978-3-030-12082-5_55-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000530_978-3-030-12082-5_55-Figure1-1.png", + "caption": "Fig. 1 Geometry of macromodel (cross section shown, dimensions in mm). Sg\u2212p\u2014contact surface of gyro unit-platform", + "texts": [ + " This, apparently, can explain the noticeable change in the measured thermal contact conductance as a function of the deformation rate in the hot forging of metal, described in [33]. Faster loading forms a different temperature field of the die and the billet and, consequently, changes the thermal contact conductance. Another factor leading ultimately to transient thermal contact conductance is thermal expansion. These phenomena can have a significant impact specifically on macroscale constructions. In the right-handed Cartesian reference system, let us consider the hypothetical construction of a gyro unit attached to a platform (Fig. 1). The model of the gyro unit is a solid cylinder with a flange, made of AISI 1020 steel. Inside the gyro unit, we assume the presence of a thermostatic systemmaintaining on a constant level the temperature of some internal part of the device. At the same time, the temperature of the device body is not constant and is determined, among other things, by interaction with the platform. The model does not consider the internal parts and the thermostatic system as separate bodies, but it takes into account that this system generates heat with a power of 20 W, corresponding to a specific heat generation rate q 94,832 W/m3 for the given volume of the gyro unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000132_j.jfluidstructs.2019.02.019-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000132_j.jfluidstructs.2019.02.019-Figure10-1.png", + "caption": "Fig. 10. Cross section of a filament in a soap film.", + "texts": [ + " (3) as shown in Fig. 9. As shown in the figure, for a certain diameter of the filaments, the frequency shows a proportion of (\u03c1fU2/\u03c1sdL)1/2. This is attributed to Argentina\u2019s theory that is based on a linear analysis such that the tolerance between real situation and theory increases when the flapping frequency increases. However, this proportionality increases with an increase in the diameter of the filaments, which seems unexpected in previous 3D experiments. Then we try to explain this phenomenon. Fig. 10 shows the cross section of a filament in a soap film. As aforementioned, the film thickness of the parallel part is far less than the diameter of the filament, so that the filament is not completely immersed in the film. In Argentina\u2019s theory, the typical flapping frequency \u03c9 is given by balancing plate inertia \u03c1sd\u03c92\u03c4 with the aerodynamic force \u03c1fU2 (\u03c4/L), where \u03c4 is the ratio of filament\u2019s lateral displacement Y to the filament length L. Here in two-dimensional fluids, the aerodynamic force decreases to \u03c1fU2 (\u03c4/L) \u00d7 \u03b4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000897_icra.2019.8793651-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000897_icra.2019.8793651-Figure1-1.png", + "caption": "Fig. 1. A compass gait model consists of two stiff legs with two degrees of freedom in a sagittal plane. The counterclockwise direction of \u03b81 and \u03b82 is defined as positive. The forward direction of xst and xsw is defined as positive.", + "texts": [ + " Effectiveness of the balance map analysis as a measure of walking balance was examined by computer simulations of unperturbed and perturbed walking during a swing phase. In the balance map analysis, it is assumed that the walking pattern is governed by the gravity and inertia. In other words, we assume a ballistic walking model [10], [13] for determining the walking balance. To simplify the bipedal mechanism, we consider a compass gait model consisting of 978-1-5386-6027-0/19/$31.00 \u00a92019 IEEE 5260 two stiff legs as shown in Fig. 1. The dynamics equation of the compass gait model is described with generalized coordinates \u03b8 = (\u03b81, \u03b82) T as follows: M(\u03b8)\u03b8\u0308 + V (\u03b8, \u03b8\u0307) +G(\u03b8) = 0, (1) M = ( (m1 +m2)l 2 1 \u2212m2l1l2C12 \u2212m2l1l2C12 m2l 2 2 ) , V = ( m2l1l2\u03b8\u0307 2 2S12 \u2212m2l1l2\u03b8\u0307 2 1S12 ) , G = ( \u2212(m1 +m2)gl1S1 m2gl2S2 ) . C12, S12, S1, and S2 denote cos(\u03b81\u2212\u03b82), sin(\u03b81\u2212\u03b82), sin \u03b81, and sin \u03b82, respectively. In addition, m1 denotes the mass of the body and stance leg, and m2 denotes the mass of the swing leg. Assuming that \u03b81 \u2248 0 and \u03b82 \u2248 0, the following linear differential equation is obtained: X\u0308 = AX (2) A = ( \u03a92 1 K (1\u2212K)\u03a92 1 K \u2212\u03a92 2 K \u2212\u03a92 2 K ) , (3) where X = (xst, xsw)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002791_s11431-020-1569-6-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002791_s11431-020-1569-6-Figure1-1.png", + "caption": "Figure 1 Mechanical and stability analyses of an SLE under axial loads. (a) Mechanical analysis of an SLE; (b) stability analysis of an SLE.", + "texts": [ + " Finally, in order to demonstrate the method presented in this work, the application of scissor deployable structures consisting of different number of units (3, 5 and 8 units) is analyzed separately and discussed comprehensively. The linear array deployable structure consists of many substructures with the same geometrical properties, and each substructure consists of a pair of bars which are connected by a pivot at the midpoint o. Due to the symmetry and consistency of the structure, it is necessary to analyze buckling of an SLE before studying the stability of the whole structure. In Figure 1(a), the AB and CD bars can rotate with each other around the pivot, and the end points A and C of these two bars are hinged on the fixed surface. Also, these two bars have the same length, and the end points B and D bear the horizontal axial loads, P. It should be noted that the instability of deployable structure under symmetrical load is studied here. For the instability under asymmetrical load, the research process is similar. Moreover, assuming these two bars are homogeneous straight rods and the friction at the hinge is not considered, the internal force generated in the joint and static derivation process can be expressed as F P F P F P F F F P F P F P F P F = , =2 tan , = , = 0, = 0, = 2 tan , = , = 2 tan , = , = 0, (1) Ax Ay Bx By ox oy Cx Cy Dx Dy where FAx, FAy, FBx, FBy, Fox, Foy are the corresponding internal forces in bar AB (similarly, bar CD) at the hinge, and \u03b3 is the deployed angle from the horizontal position to the unit member. The SLE connected by bars AB and CD in the hinge o is subjected to horizontal axial loads, so it has a larger deformation around the midpoint o depicted in Figure 1(b). Moreover, because of the symmetry of the SLE configuration and force depicted in Figure 1(a), the stability of the whole unit can be represented by a bar. Therefore, bar AB is used to analyze the buckling load of the SLE at the time of instability. A-xy is the global coordinate system and the bodyfixed coordinates o1-x1y1 and o2-x2y2 are attached at the points A and B of the bar AB. In the AO section, we assume that the bending deformation occurs in a main inertial plane of the bar, and the direction along x1 is positive in the body-fixed coordinate system o1x1y1. The bending stiffness is EI", + " (10) shows that the displacement, \u03b4, normal to the axis of the bar tends to infinity when the variable \u03b1L increases to \u03c0 or \u03b2L increases to \u03c0/2, which indicates that the deployable structure no longer has the ability to bear the load at this time, and any small load will cause this structure to collapse. Therefore, when \u03b1L=\u03c0 or \u03b2L=\u03c0/2, the load obtained is what we need, namely the buckling load. In this situation, it can be noted that no matter how small the load for an SLE is, the structure will be destroyed. Correspondingly, the obtained buckling load can be written as P EI L P EI L = (cos + 2tan sin ) , = 4 cos . (11) 1 2 2 2 2 2 In eq. (11), P1 represents the buckling load of AO and CO sections presented in Figure 1(b), and P2 indicates the buckling load of BO and DO sections. It should be highlighted that the smaller of P1 and P2 values is chosen as the critical load of the SLE. When the mechanical analysis and stability analysis of an SLE are completed, the next step is to focus on the buckling conditions of the deployable structure according to the contribution of each SLE. Deployable structure consists of units which are composed of pairs of bars connected at a joint that allows a compact and deployable configuration" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003215_012013-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003215_012013-Figure1-1.png", + "caption": "Figure 1. Cardan shaft of the driveline for Automobile", + "texts": [ + " To reduce power loss in the powertrain, composite materials can be used to design Cardan shafts with less weight while maintaining the required strength and stiffness [2]. A researched by performing FEA analysis on two different composite materials are E-GLASS and E-CARBON reducing Von-Mises stress and weight compared to structural steel. This can be concluded that the composite material can be used in the Cardan shaft [3]. 10th TSME-International Conference on Mechanical Engineering (TSME-ICoME 2019) IOP Conf. Series: Materials Science and Engineering 886 (2020) 012013 IOP Publishing doi:10.1088/1757-899X/886/1/012013 Figure 1 shows the Cardan shaft for the automobile. The Cardan shaft is used to connect the gearbox to the different rear axle, and to transfer power from the engine to differential automobile through universal joints assembly. Cardan shaft has to withstand high rotational speeds and has to change its length while transmitting the torque required by the vehicle. The second method used to identify the motion of the universal joints which is proportional to the angle between the driving and driven shafts. It also affects system vibrations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001683_icems.2019.8921881-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001683_icems.2019.8921881-Figure6-1.png", + "caption": "Fig. 6. Design of rotor pole. (a) One-side groove; (b) symmetry grooves.", + "texts": [ + "d Besides, the average torque appears to decrease gradually compared with the initial SRM. For doubly salient motor, the fringing flux effect and local magnetic saturation before the overlap angle are common problems, which can be alleviated by changing the shape of the salient pole. Thus, some scholars have proposed a new rotor pole shape with groove in the forward rotating direction, but it is impossible to suppress the torque ripple when SRM is reversely rotating. Therefore, a rotor pole with symmetrical grooves on both sides is studied, as seen in Fig. 6. Since the opening width of the groove is depended on the range of fringing flux, the shape of the groove is controlled by the groove angle \u03b8 and groove length L . According to the initial parameters, \u03b8 will vary from 0 to 90\u00b0 and L will grow from 0 to 6 mm. In order to verify the superiority of the symmetric rotor pole, the FE models of rotor poles with grooves on one side and with bilateral grooves are set up, which keep other parameters as initial and choose 30 ,\u03b8 = \u00b0 3 mm .L = The comparison of the two rotor shapes is presented in Table III" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002705_kem.841.144-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002705_kem.841.144-Figure2-1.png", + "caption": "Fig. 2 Macroscopic view of the broken spur gear", + "texts": [ + " (#541197814, University of Melbourne, Melbourne, Australia-26/07/20,01:08:57) Material. This study examined the spur gears of an incapacitated hand tractor after operating for only 40 hours, classified as premature failure. This research was conducted using empirical methods and the finite element method (FEM). The subject of this study is illustrated in Fig. 1. Visual Inspection. Initially, a detailed visual inspection of the hand tractor was conducted to assess the general quality of the gears and to identify all relevant fracture features, as displayed in Fig. 2. Chemical Composition Analysis. Chemical analysis was also carried out to determine the chemical composition of the gears so that it could be classified correctly into the proper standard groups, such as ASM or AISI. Hardness Testing. Hardness testing was performed at 9 points on the outer and inner surface of the specimen. The specimen was initially separated by cutting in half to measure the hardness of the inner surface. The test was performed employing the Rockwell method the using HRB unit [7]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001412_mees.2019.8896368-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001412_mees.2019.8896368-Figure2-1.png", + "caption": "Fig. 2. Angular characteristic of the power of synchronous electric drive.", + "texts": [ + "3 \u2013 block of formation of the corresponding control; 13 \u2013 direct forced excitement; 14 \u2013 Executive excitation transmission system. In the synchronization mode of the electromotive force, the force (Eq) must satisfy the following equation: )( 000 GGUqq UUKEE where, Eq0 \u2013 default value of EMF; UG0 \u2013 default value of voltage; K0U \u2013 amplification coefficient of ASFE according to motor voltage. The dynamics of the electromagnetic and electromechanical processes course correspond to the external angular characteristic, shown on Fig. 2. The steady-state mode corresponds to the point a with coordinates P0, \u03b80 at the corresponding values of the EMF used as initial value. During investigation, it is necessary to increase or decrease the active power of the synchronous machine, depending on the mode of loading. Outside the gn - gb angular characteristic coincides with the marginal internal characteristics that may correspond to Eqmin and Eqmax which are used as the limit values of the synchronous EMF. An outstanding feature of the angular characteristic is the shift of t PM to the right according to the extreme values, the limits m\u2019 and m, correspond to the extreme power values" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003131_j.optcom.2020.126271-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003131_j.optcom.2020.126271-Figure1-1.png", + "caption": "Fig. 1. Schematic 3D and cross-sectional structure of the proposed telescope objective: (a) Schematic 3D of the proposed telescope objective. (b) Sectional view of the liquid prism.", + "texts": [ + " In particular, there is a transparent partition sheet with the same density as these two liquids between L\u2013L interface, which help to obtain perfect flat interface. Driving by electrowetting effect, the largest tilted angle of the L\u2013L interface is \u223c52.4\u25e6 and the light beam steering angle is \u223c10.4\u25e6. The tunable FOV of the proposed telescope objective with this liquid prism is \u223c5.8\u25e6. The response time is about 480 ms. Comparing with the conventional telescopic objective, the proposed device is very compact and easy to operate. The proposed telescope system is shown in Fig. 1(a). It consists of three main parts: telescopic objective, liquid prism and CMOS. The proposed liquid prism is the essential part, and its schematic 3D and sectional view are shown in Fig. 1(b), respectively. This device is a cubic chamber composed of four electrodes coated with hydrophobic insulator respectively. It is filled two immiscible liquids with equal densities but different refractive indices. One of the liquids is electrically conductive, the other is insulative. Especially, the conductive liquid contacts the substrate electrode. Different from the conventional liquid prism, a transparent partition sheet is placed between the L\u2013L interface. Therefore, a flat interface is formed in the cubic chamber" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000755_b978-0-08-102211-5.00004-8-Figure4.32-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000755_b978-0-08-102211-5.00004-8-Figure4.32-1.png", + "caption": "Figure 4.32 Auxetic plied yarn at initial state: (A) side view and (B) after spiral expansion [27].", + "texts": [ + " The cross section of the APY after the two stiff yarns get in contact with each other and the movement path of the stiff yarn during the first stage of stretch are shown in Fig. 4.31B and C, respectively. Like the previous analysis, the variation of the effective diameter of the APY (H) in this model is also utilised to calculate the Poisson\u2019s ratio. To do that, the key is still to calculate the distance between the central point of the APY to the surface point of stiff yarn or soft yarn. The first step is to determine the initial tilt angle \u03b40. Fig. 4.32 shows the side view of the yarn and the expansion of both the stiff and soft yarns in one turn of the APY. From Figs 4.31A and 4.32, the following equations can be derived: ystiff 5 Sd0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u00f0d20=4\u00de2 x2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi S2d02L20 p =2\u03c0 cos\u03b40 2 r L0 1 sin\u03b40 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi S2d0 2 L20 p 2\u03c0 (4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure1.12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure1.12-1.png", + "caption": "Fig. 1.12 Fuel pump used in Delphi Econolyst case study (Images are taken from Reeves 2013)", + "texts": [ + " In the Delphi study, two materials; the powder form of the cast aluminum alloy A380.0 used in the original component and a stainless steel (316L); are studied using a selective laser melting process. The study not only analyzes the energy associated with manufacturing the conventional and AM components, but also quantifies the in-use and post-use energy, encompassing a cradle-to-grave\u2013based analysis. The resulting AM and conventional manufacturing designs are presented in Fig. 1.10. Although the designs of the AM and the CM components in Fig. 1.12 are visually very different, there is no geometric difference in fluid path connection or pump drive locations between the two components. Besides the opportunity for improved flow 16 J. Williams et al. path and reduced secondary machining operations, the AM component also had considerable weight savings compared with the conventional component, even if made from ~2X denser 316 SS. These weight savings not only translate to in-use benefits but also are realized at the raw material acquisition, transport and production stages, as less material is required overall" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003455_1350650120954943-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003455_1350650120954943-Figure2-1.png", + "caption": "Figure 2. Kinematics diagram of ABB IRB 6620.", + "texts": [ + " Articulated manipulators are prevailing and popular in industrial scenarios, and thus their motion control problem is of great significance. Take ABB IRB 6620, a typical six axes commercial industrial manipulator as an example for the case study, we can obtain insights about the effects of friction on dynamics and motion control, taking the complex environmental factors such as load and temperature into consideration. Kinematics of ABB IRB 6620 is first analyzed, and its modified D-H parameters are obtained from the CAD model provided at ABB\u2019s official website, which is shown in Figure 1. Figure 2 demonstrates the kinematics diagram of all six joints in ABB IRB 6620 manipulator adopting a modified D-H style, and its detailed kinematics parameters are displayed in Table 1, where the last column shows the initial position. With inertial parameters such as the center of mass and mass of links taken from the CAD model, we could obtain the dynamic model without friction first, by adopting iterative Newton-Euler dynamics algorithm2: s \u00bc M\u00f0H\u00de \u20acH \u00fe C\u00f0H; _H\u00de _H \u00fe G\u00f0H\u00de (1) where \u20acH; _H;H are joint acceleration, velocity and position respectively, s is driving torque applied by motor, M\u00f0H\u00de and C\u00f0H\u00de stand for inertial and centrifugal and Coriolis matrix respectively and G\u00f0H\u00de represents the gravitational torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002299_rteict42901.2018.9012428-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002299_rteict42901.2018.9012428-Figure10-1.png", + "caption": "Fig. 10. Magnetic Parameters of 50% Short LSPMSM. (a) Radial Flux Density. (b) 2-D Flux Density Distribution. (c) Spatial FFT of Radial Flux Density distribution.", + "texts": [ + " 7(c), which shows the introduction of noise content in the FFT spectrum due to asymmetry created by short-circuit fault compared to healthy condition. B. Parameters of 50% short LSPMSM: The 50% shorted phase current shows a further increase in magnitude of nearly 9A peak and decrease in induced voltage as depicted in Fig. 8. After the incorporation of 50% inter-turn fault, the speed and torque ripples increases significantly compared to 25% short-circuit condition as depicted in Fig. 9(a) and Fig. 9(b). From Fig. 10(a), the augmented un-symmetry in the radial flux density of the faulty machine as compared to 25% short can be noticed. The saturation of flux density distribution in some parts of the machine is encircled in Fig. 10(b). The spatial FFT of radial flux density is shown in Fig. 10(c), which shows noise content in the spectrum due to asymmetry in the radial airgap fields. The results of co-simulation of healthy and faulty LSPMSM indicate that as the intensity of stator turn-to-turn short-circuit fault increases, the magnitude of stator current in the corresponding shorted phase increases proportionately. Also, oscillations can be noted in the moving speed of faulty machine leading to increased torque ripple. The radial flux density distribution becomes more asymmetrical and distorted as the fault condition degenerates. The flux distortion can be observed from the 2-D geometrical view of the faulty machines in the regions encircled black in Fig. 7(b) and Fig. 10(b). Also the spatial FFT of radial flux density shows that the flux density of the faulty machine is more polluted with sub-harmonics when compared to that of the healthy machine. Parameters which show measurable changes when there is a stator turn-to-turn SC fault are Stator Current, Machine Induced Voltage and Radial air-gap Flux Distribution. Among these, stator current is the simplest quantity which is easily accessible and directly available for analysis. Hence MCSA can be a suitable fault detection method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000768_s11340-019-00524-0-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000768_s11340-019-00524-0-Figure1-1.png", + "caption": "Fig. 1 Tension specimen design for high-speed X-ray imaging. Dimensions are in mm", + "texts": [ + " Therefore, specimens were limited to this approximate thickness. In an attempt to adhere to existing standards for dynamic experimentation and the rigorous requirements for experimentation at the APS, specimens were designed to reflect a scaled and thinned version of the ASTM D638 Type V specimen. The specimen design reflected an approximate scaling of 31% of the standard, while simultaneously being thinned to a nominal thickness of 100 \u03bcm. The engineering drawing for the specimens is shown in Fig. 1. Specimens were cut from a block of 316 L material produced by a 3D Systems ProX 200 (3D Systems \u00a9) direct metal printer. A 1070 nm continuous wave laser with a spot size of approximately 100 \u03bcm was used to melt the powder stock with a power of 103 W, and with a scanning velocity of 1400 mm/s. The \u201chexagons\u201d scan strategy was used to provide a laser path [31]; layer thickness was 30 \u03bcm, and the oxygen content in the chamber was approximately 1000 ppm. The average particle diameter of the powder stock was 17" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003838_s11071-020-06016-4-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003838_s11071-020-06016-4-Figure1-1.png", + "caption": "Fig. 1 Disk on inclined support", + "texts": [ + " It is important that, in contrast to the other quantities, R cannot be predefined uniquely: we should first specify contact conditions. Clearly, in the absence of the contact R \u00bc 0, but in case of contact the possibilities of rolling and sliding (both directions) are to be considered in turn, not to mention a jump. This makes dynamics of seemingly simple problem rather sophisticated and sometimes even paradoxical. 2.1 Mathematical description To develop a mathematical description, introduce an inertial frame with the axis OX within the plane and OY directed outside (Fig. 1). Let \u00f0x; y\u00de be the coordinates of pointC and h be the angle betweenOX andCG, then the variables in (2.1) have the following coordinates: v \u00bc _x a _h sin h; _y\u00fe a _h cos h ; x \u00bc _h; q \u00bc a cos h; y\u00fe a sin h\u00f0 \u00de; P \u00bc mg\u00f0sin c; cos c\u00de System (2.1) becomes m d2 dt2 x\u00fe a cos h\u00f0 \u00de \u00bc mg sin c\u00fe T ; m d2 dt2 y\u00fe a sin h\u00f0 \u00de \u00bc mg cos c\u00fe N; mk2 d2h dt2 \u00bc y\u00fe a sin h\u00f0 \u00deT \u00fe Nr cos h \u00f02:2\u00de where N and T are the normal reaction and friction force. In the case of lasting contact y r. Introduce new dimensionless variables by the formulas T \u00bc T= mg cos c\u00f0 \u00de; N \u00bc N= mg cos c\u00f0 \u00de; x \u00bc x=r; y \u00bc y=r; a \u00bc a=r 1; k \u00bc k=r \u00f02:3\u00de as well as a new independent variable s \u00bc t ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi g cos c=r p , then system (2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure9-1.png", + "caption": "Figure 9. Pressures presentation", + "texts": [], + "surrounding_texts": [ + "Radial Load Effect Pressure Distribution for Textured Plain Bearing To\ufeff demonstrate\ufeff the\ufeff effect\ufeff of\ufeff the\ufeff radial\ufeff load\ufeff on\ufeff the\ufeff operating\ufeff performance\ufeff of\ufeff the\ufeff non-textured\ufeff hydrodynamic\ufeffplain\ufeffbearing,\ufeffsuch\ufeffas\ufeffthe\ufeffpressure\ufeffdistribution\ufeffand\ufeffthe\ufefffluid\ufeffflow\ufeffvelocity\ufeffwithin\ufeffthe\ufeff plain\ufeffbearing,\ufeffa\ufeffradial\ufeffload\ufeffvariation\ufeffis\ufeffperformed\ufeff(W1\ufeff=\ufeff2000N,\ufeffW2\ufeff=\ufeff6000N\ufeffand\ufeffW3\ufeff=\ufeff10000N).\ufeff The\ufeffinitial\ufeffoperating\ufeffconditions\ufeffof\ufeffthe\ufeffbearing\ufeffare\ufeffwith\ufeffa\ufeffsupply\ufefftemperature\ufeffTa\ufeff=\ufeff40\ufeff\u00b0C,\ufeffsupply\ufeff pressure\ufeffPa\ufeff=\ufeff0.04\ufeffMPa\ufeffand\ufeffthe\ufeffspeed\ufeffof\ufeffrotation\ufeffof\ufeffthe\ufeffshaft\ufeffequal\ufeffto\ufeffN\ufeff=\ufeff6000\ufeffrpm.", + "Figure\ufeff10\ufeffshows\ufeffthe\ufeffdistribution\ufeffof\ufeffthe\ufeffpressure\ufeffalong\ufeffthe\ufeffmedian\ufeffplane\ufeffof\ufeffthe\ufeffplain\ufeffbearing\ufefffor\ufeff different\ufeffradial\ufeffloads.\ufeffThe\ufeffsignificant\ufeffpressures\ufeffare\ufeffobtained\ufefffor\ufeffa\ufefftextured\ufeffplain\ufeffbearing\ufeffunder\ufeffa\ufeffradial\ufeff load\ufeffof\ufeff10000N.\ufeffAs\ufeffwell\ufeffas\ufeffthe\ufeffincrease\ufeffof\ufeffthe\ufeffradial\ufeffload\ufeffleads\ufeffto\ufeffan\ufeffincrease\ufeffof\ufeffthe\ufeffhydrodynamic\ufeff pressure.\ufeffThis\ufeffincrease\ufeffreaches\ufeff75\ufeffpercent\ufefffor\ufeffthe\ufefftextured\ufeffplain\ufeffbearing\ufeffcase.\nThe\ufeffpressure\ufeffdistribution\ufeffaccording\ufeffto\ufeffangular\ufeffposition\ufefffor\ufeffplain\ufeffbearing\ufefftextured\ufeffand\ufeffnot\ufefftextured\ufeff for\ufeffradial\ufeffload\ufeffof\ufeff10\ufeff000N\ufeffand\ufeffrotational\ufeffvelocity\ufeffof\ufeff9000rpm\ufeffis\ufeffpresented\ufeffrespectively\ufeffin\ufeffFigure\ufeff 10a\ufeffand\ufeffFigure\ufeff11.\ufeffThe\ufeffcurve\ufeffclearly\ufeffshows\ufeffthat\ufeffthe\ufeffpressure\ufeffdistribution\ufeffalong\ufeffthe\ufeffmedian\ufeffplane\ufeffof\ufeff the\ufeffplain\ufeffbearing\ufeffis\ufeffdifferent\ufefffor\ufeffthe\ufeffcase\ufeffof\ufeffa\ufeffplain\ufeffbearing\ufeffwithout\ufefftexture\ufeffand\ufeffa\ufeffplain\ufeffbearing\ufeffwith\ufeff textured\ufeffsurface.\ufeffthe\ufeffdifference\ufeffis\ufeffestimated\ufeff75\ufeffpercent.\nDisplacement Distribution for Textured Bearing The\ufeffelastic\ufeffdisplacement\ufeffof\ufeffthe\ufeffbearing\ufeffinner\ufeffface\ufeffis\ufeffdue\ufeffto\ufeffthe\ufeffdeformation\ufeffof\ufeffthe\ufeffbearing\ufeffgenerated\ufeff by\ufeffthe\ufeffhydrodynamic\ufeffpressure\ufefffield,\ufeffis\ufeffshown\ufeffin\ufeffFigure\ufeff12\ufefffor\ufeffa\ufeffload\ufeff2000\ufeffN\ufeffand\ufeffthe\ufeffbearing\ufeffoperates\ufeff at\ufeffa\ufeffrotational\ufeffvelocity\ufeffof\ufeff6000\ufeffrpm.\ufeffThe\ufeffmaximum\ufeffdeformation\ufeffis\ufeffnoted\ufeffat\ufeffthe\ufefflevel\ufeffof\ufeffthe\ufefflower\ufeff generatrix\ufeffof\ufeffthe\ufeffbushing.\nFigure\ufeff13\ufeffillustrates\ufeffthe\ufeffdisplacement\ufeffdistribution\ufeffin\ufeffthe\ufeffcircumferential\ufeffdirection\ufeffof\ufeffthe\ufeffplain\ufeff bearing\ufeffas\ufeffa\ufefffunction\ufeffof\ufeffthe\ufeffradial\ufeffload\ufeffand\ufefffor\ufeffthe\ufeffrotation\ufeffvelocity\ufeffof\ufeff6000\ufeffrpm.\ufeffThe\ufeffcurve\ufeffshows\ufeff that\ufeffthe\ufeffdeformation\ufeffhas\ufeffa\ufefflarge\ufeffincrease\ufeffwith\ufeffthe\ufeffrise\ufeffof\ufeffthe\ufeffload,\ufeffthis\ufeffincrease\ufeffreaches\ufeff0.35\u03bcm\ufefffor\ufeff a\ufefftextured\ufeffsurface\ufeffplain\ufeffbearing;\ufeffon\ufeffthe\ufeffother\ufeffhand,\ufeffreaches\ufeffthem\ufeff0,16\u00b5m\ufefffor\ufeffa\ufeffnon-textured\ufeffplain\ufeff bearing,\ufefffor\ufeffthe\ufeffcase\ufeffof\ufeffa\ufeffplain\ufeffbearing\ufeffsubmitted\ufeffa\ufeffloading\ufeff10000N.\nFigure\ufeff14\ufeffshows\ufeffthe\ufeffdisplacement\ufeffalong\ufeffthe\ufeffcircumference\ufeffdirection\ufeffof\ufeffthe\ufeffpad.\ufeffDisplacement\ufeff is\ufeffvery\ufeffimportant\ufefffor\ufeffthe\ufeffcase\ufeffof\ufeffa\ufefftextured\ufeffbearing\ufeffand\ufeffvery\ufeffloaded,\ufeffit\ufeffforms\ufeffa\ufeffkind\ufeffof\ufeffa\ufeffsignificant\ufeff hollow\ufeffat\ufeffthe\ufefflevel\ufeffof\ufeffits\ufefflower\ufeffbearing\ufeffgeneratrix\ufeffcompared\ufeffto\ufeffthat\ufeffnoted\ufefffor\ufeffa\ufeffnon-textured\ufeffbearing.\ufeff" + ] + }, + { + "image_filename": "designv11_80_0003451_14763141.2020.1805507-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003451_14763141.2020.1805507-Figure2-1.png", + "caption": "Figure 2. Comparison showing constant cut-off frequency (fc) of conventional filter and time-varying fc of the fractional Fourier filter. X1 = cut-off frequency of swing phase, W = width of impact, H = height of impact (maximum cut-off frequency), ti = time of impact centre.", + "texts": [ + " The details of each CF are outlined in Table 1. For the FrFF condition, marker trajectories were exported to Matlab (2017b, Natick, USA) where custom-written scripts performed filtering (Georgakis & Subramaniam, 2009). The FrFF employs a triangular filter boundary that raises the cut-off frequency to retain high-frequency motion content from physical sources (i.e., deceleration of lower leg) during impact (Georgakis & Subramaniam, 2009), and boundary parameters were determined as per Augustus et al. (2020; Figure 2). Parameter X1 represents the cut-off frequency of the non-impact phase and was set as 22 Hz (determined by residual analysis), W the duration of impact (calculated as time of BCE minus time of BCS), tI the time of maximum acceleration induced by impact and H the height of impactinduced expansion of the frequency content at tI (i.e., maximum physical frequency induced during impact). Since the physical characteristics of each marker trajectory were different, parameters W and H were optimised on a trial by trial basis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001376_j.mechmachtheory.2019.103672-Figure14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001376_j.mechmachtheory.2019.103672-Figure14-1.png", + "caption": "Fig. 14. Rectifier mechanism.", + "texts": [ + " 13, depending on several parameters such as ring\u2019s inertia, epicyclic train ratio, input four-bar mechanism design, input angular velocity, etc.; therefore, a specific study is needed for each individual gearbox system (a more detailed explanation about this phenomenon can be seen in [19]). In the rectifier mechanism, the sun gear of the inertial epicyclic train has an oscillating movement which must be rectified in order to be used by a vehicle. The rectification is carried out by the rectifier mechanism (Fig. 2). This mechanism can be built in different ways; as an example, the system schematised in Fig. 14 is used for the analysis presented hereinafter. In this mechanism, when the sun of the epicyclic train is turning clockwise, positive freewheel is blocked and the power is transmitted to the output through the gears G1, G2 and G3. At the same time, negative freewheel is decoupled. In the other half period, the sun turns counter-clockwise and negative freewheel is blocked and the positive one decoupled. In this situation, the power is transmitted through gears G4 and G5 (a detailed explanation of this mechanism can be found in [2])" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001549_s11223-019-00110-8-Figure17-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001549_s11223-019-00110-8-Figure17-1.png", + "caption": "Fig. 17. Structural sketch of integrated meshing stiffness.", + "texts": [ + " Integrated Meshing Stiffness. Considering thermal strain, we divided the integrated meshing stiffness of the movable tooth pair into two parts: the contact meshing stiffness and thermal meshing stiffness. The meshing stiffness is equivalent to a \u201cspring\u201d model. Because thermal elastic deformation occurs on the basis of the initial deformation, the two \u201csprings\u201d are superimposed on each other. Therefore, the integrated meshing stiffness becomes \u201cflexible,\u201d and its structural sketch is shown in Fig. 17. Integrated meshing stiffnesses kH of the wave generator and movable tooth and kK of ring gear and movable tooth can be calculated as follows: k k k H HT HN 1 1 1 , (19) k k k K KT KN 1 1 1 . (20) Fig. 14. Thermal strain cloud atlas of the movable tooth. Fig. 13 Fig. 14 Fig. 15 Fig. 16 Fig. 15. Thermal meshing stiffness of the wave generator and movable tooth. Fig. 16. Thermal meshing stiffness of the movable tooth and ring gear. The integrated meshing stiffness data of the meshing pairs are plotted in Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001333_aim.2019.8868416-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001333_aim.2019.8868416-Figure1-1.png", + "caption": "Fig. 1. Setup in a robot-assisted laparoscopic hysterectomy. (a) configurations of the robot, the laparoscope, and the laparoscopic tool; (b) relationship between the coordinate frames", + "texts": [ + "00 \u00a92019 IEEE 169 of the uterus nor the laparoscopic tool is visible from the laparoscope once calibrated, (3) changing of laparoscopic tool can be performed as usual; as the sensor module is integrated to the trocar, it will not affect the laparoscopic tool, (4) sensors are outside the patient\u2019s body, which makes its sterilization requirements less strict, and (5) by using the system, the user can use his/her hands to control the robot without releasing them from the laparoscopic tool. Experiments are conducted to verify our proposed approach. It is demonstrated that maximizing the distance between the fundus of the uterus and the laparoscopic tool tip can be achieved in an automatic manner with our proposed approach. In the section, we present the detailed working principle of our proposed method. Fig. 1(a) illustrates the setup during a robot-assisted laparoscopic hysterectomy proposed in [18]. Incisions are opened on the patient\u2019s abdomen for the access of the laparoscope and the laparoscopic tools. From the end of the operating table, a surgical tool (i.e. uterine manipulator) is inserted to the patient\u2019s body for uterus manipulation. In [18], a prototype is developed to assist in this uterus manipulation task. In this paper, we assume the captioned prototype (or similar) is used and uterus manipulation is performed by the prototype about a virtual pivot point located at the cervix", + " the base frame of the robot is fixed) \u2022 Yaw and pitch motions of the trocar about the incision point is allowed while insertion along the incision is prohibited \u2022 Insertion of the laparoscopic tool is allowed along the axial direction of the trocar \u2022 Yaw and pitch motions of laparoscopic tool reflect to the trocar \u2022 Total length of the laparoscopic tool is known \u2022 Distance between the incision point and the range sensor of the pose sensing trocar is known \u2022 The position and orientation of the laparoscopic camera is fixed during the calibration period \u2022 The position of the tip of the laparoscopic tool and the robot\u2019s end-effector in the image space can be obtained Assume the virtual pivot point of the uterus positioning robot is fixed and the point of incision is also fixed. And, the relationship between the base frames of the uterus positioning robot, the laparoscope, and the trocar is as illustrated in Fig. 1(b). Based on the pinhole camera model [19] (see Eq. 1), the extrinsic parameters which describe the relationship between the camera frame {Fc} and the world coordinate frame can be calibrated with sufficient observations. Here, we assume there are two world coordinate frames, the robot\u2019s frame {Fr} and the trocar\u2019s frame {Ft}. Assume the camera is fixed at the moment, with reference to Fig. 1(b), the three coordinate frames form a closed loop. Thus, the homogeneous transformation matrix r t T which relates {Ft} to {Fr} can be obtained by Eq. 2. Note that since the virtual pivot point of the robot (or, the base frame of the robot) and the incision point are fixed, {Ft} and {Fr} are fixed; thus r t T would not change even if the camera moves. s [ u v 1 ]T = K[R|t] [ X Y Z 1 ]T (1) where s is the scaling factor, [ u v ]T is the position in the image space, K is the intrinsic parameters, [R|t] is the extrinsic parameters, and [ X Y Z ]T is the position in the world coordinate frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003914_gucon48875.2020.9231153-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003914_gucon48875.2020.9231153-Figure1-1.png", + "caption": "Fig. 1. Configuration of quadcopter from earth to body fixed frame", + "texts": [ + "00 \u00a92020 IEEE 718 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on June 04,2021 at 09:32:01 UTC from IEEE Xplore. Restrictions apply. These assumptions are considered for deriving the model [18][19]- i. The structure of quadcopter must be rigid and symmetrical. ii. The center of gravity and origin of Body fixed frame must coincide. iii. Mass of the quadcopter and moment of inertia is taken constant. The four inputs of the quadcopter are basically the four fixed-pitch rotors. The thrust provided by each propeller of quadcopter is shown in figure 1. Here considering a body frame as B and Earth frame as E. Using parameterization of Euler angles, the orientation of airframe in space is given by RT from body to earth. The Newton-Euler formalism is used in deriving the dynamic model of the quadcopter as shown in [18]-[20]. The relation between the generated thrust and the propeller\u2019s angular velocity is defined by equation (1) and between thrust and torque of the propeller by equation (2)- (1) (2) Where represents the thrust generated and represent torque generated by the nth propeller respectively and represents the angular velocity of nth propeller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002274_j.promfg.2020.02.008-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002274_j.promfg.2020.02.008-Figure1-1.png", + "caption": "Fig. 1. Schematic of direct energy deposition.", + "texts": [ + " Furthermore, the heat-input conditions are also evaluated with a heat conductivity simulation to keep the cooling rate constant theoretically. As a result, the experimental results show that even quality in Vickers hardness and metal structure can be ensured by employing cooling system and preheating process on the baseplate at the same time. Directed energy deposition (DED) is one of the AM available for metals. According to the definition of DED decided by ASTM, other heat sources like electric arc are also available, and wire material is also used. Figure 1 shows a schematics of laser-based powder DED, which is used in this study. Irradiating the laser beam on the baseplate to form a melt pool and supplying the material powder with carrier gas, the deposit contentiously gets large by partially repeating solidifying and melting the supplied powder. Furthermore, the shield gas is also supplied to protect the laser system from the sputter generating of melt pool. Generally, inert gas like Argon is used for the carrier and shield gases to avoid oxidation on the deposit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003854_s11465-020-0597-z-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003854_s11465-020-0597-z-Figure1-1.png", + "caption": "Fig. 1 Structure of the four-side PPMLM drive press: (a) Magnified view and (b) front sectional view.", + "texts": [ + "cn Field-circuit coupling simulation is conducted to evaluate the system characteristics. For an experimental study, a 6 kN press prototype is manufactured, and the servo control system is implemented. Punch strokes with different speeds and trajectories are tested, and the results exhibit favorable dynamic tracking and static positioning performance. 2 PPMLM drive mechanical press 2.1 Machine structure A four-side PPMLM drive mechanical press is designed based on our previous research and shown in Fig. 1. Four symmetrical PPMLMs are arranged around the punch synchronously to balance the normal force between the primary and the secondary. The three-phase, shortprimary/long-secondary topology of PPMLM is utilized. The multi-tooth primary core modules and PMs are arranged alternately on each primary, and the end PMs are employed on both margins to compensate for the longitudinal end effect. Three-phase interval windings are wound around two adjacent primary core modules. Correspondingly, each secondary module contains only the salient pole core, which is directly installed on the punch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001429_j.mechmachtheory.2019.103674-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001429_j.mechmachtheory.2019.103674-Figure9-1.png", + "caption": "Fig. 9. Using state correction for changed mechanism to avoid unintended motion.", + "texts": [ + " Whenever a new set of parameters is sent to the PLC, a dwell trajectory for each parameter is planned with polynomials of fifth order. In order to save calculation time these polynomials are evaluated every tenth time step of the PLC. Subsequently, with a linear interpolation new set values for each time step are calculated and the structure for the mechanism is created and used for the dynamic calculations. Parameter changes can lead to motion of the handle, though the crank angle stays constant, see handle positions in Fig. 9 . Such a motion is unintended because no force was applied in direction of the trajectory. Furthermore, this may lead to high accelerations or may push the mechanism into the limit position, which again leads to high accelerations. For the compensation of this effect a state correction is introduced. In the event of parameter changes a corrected state, i.e. crank angle, is calculated. To minimize the unintended motion of the handle, the associated crank angle of the closest position of the new mechanism to the old mechanism is determined" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001126_1.5122894-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001126_1.5122894-Figure6-1.png", + "caption": "Fig. 6. Angular velocity is greater than the critical value. x\u00bc 0.62 rad s 1, hmax\u00bc 0 , and t\u00bc 1.68 s. In this case, the envelope does not contain any part of the circle.", + "texts": [ + " The parabolas that are described by the particles can be derived using vox \u00bc ax cos h; voy \u00bc ax sin h; xo \u00bc a sin h; yo \u00bc a cos h: From these, we get t \u00bc a sin hx ax cos h and y \u00bc a cos h\u00fe tan h\u00f0a sin h x\u00deg \u00f0a sin h x\u00de2 2a2x2 cos2h : (6) For the maxima of the trajectories, xmax \u00bc a2x2 cos h sin h g \u00fe a sin h; ymax \u00bc a cos h\u00fe a2x2 sin2h 2g ; where h is the angle with the y-axis for the point where the particle is ejected by the wheel. We get for the foci of the parabola Fx \u00bc a sin h a2x2 2g sin 2h; Fy \u00bc arccosh 1 2g a2x2 cos 2h: The maximum and the focus of each trajectory are depicted in Fig. 5 which corresponds to x xcritical0 . These figures also show Campbell\u2019s envelope, which depicts the position of the particles for all possible parabolas at time t. As can be seen, this envelope changes with time. In this paper, we presented two types of critical values for envelopes. In the first case, the critical value was the angle of the point of release of a particle with the vertical y-axis. This critical angle depends on the shape, where the initial points Ai are located and the positions of mirrors as a fraction of the initial height" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000810_usys.2018.8779207-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000810_usys.2018.8779207-Figure1-1.png", + "caption": "Fig. 1: Guidance geometry", + "texts": [ + " M = [m11, 0, 0; 0,m22, 0; 0, 0,m33], D(\u03bdi) = [d11, 0, 0; 0, d22, 0; 0, 0, d33]. C(\u03bdi) = [0, 0,\u2212m22vi; 0, 0,m11ui;m22vi,\u2212m11ui, 0] Assumption 1 :To simplify the algorithm design process, the gravity satisfy the condition that G(\u03b7i) = 03\u00d71. We can rewrite the ith AUV\u2019s dynamics through Eq.(1) as u\u0307i = (m22viri \u2212 d11ui + w1i + Fi) /m11 = \u03c61i + Fi/m11 v\u0307i = \u2212(m11uiri \u2212 d22vi + w2i)/m22 = \u03c62i r\u0307i = ((m11 \u2212m22)uivi \u2212 d33ri + w3i + \u03c41i) m33 = \u03c65i + \u03c41i m33 (2) where w1i, w2i, w3i are the three elements of the vector wi which denotes the ith AUVs external disturbance. In Fig. 1, engagement geometry of the ith AUV attack a target is shown In Fig.1, OXY is the inertial coordinate frame. Vt denotes the velocity of the target. Ri and Vmi represent the range-togo and velocity of the ith AUV, \u03a8i is the deflection angle of the trajectory. qi is the Line of Sight (LOS) angle. Then, the guidance kinematics model of ith AUV can be written as R\u0307i = Vt cos pti \u2212 Vmi cos pi Riq\u0307i = Vmi sin pi \u2212 Vt sin pti qti = qi = pi + \u03a8i = arctan ( yt \u2212 yi xt \u2212 xi ) (3) where xt and yt are the position of the moving target. pti is the prepositive angle. Firstly, we should employ the derivative on both sides of the first equation in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001355_s1068798x19100095-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001355_s1068798x19100095-Figure1-1.png", + "caption": "Fig. 1. Kinematics of the drive for a rolling-mill stand: (1) triple conical gear; (2) cylindrical gear; (3, 4) double conical gears; (5\u20137) rolling shafts; M, external torque.", + "texts": [ + "3103/S1068798X19100095 In the design of piecewise rolling mills for seamless pipe, careful attention is paid to overloading and dynamic decrease in speed of the drive motors, since the rolling torque depends directly on the roller speed and so slight decrease in motor speed may greatly lower the load in transient conditions. We investigate the dynamic loads in the drive of the rolling stand at the TPA 30-102 pipe-rolling system with a continuous straightening mill (Pervouralsk pipe plant). This system includes a 24-stand rolling mill with three-roller stands. Each stand has three inputs and a differential group drive. A kinematic diagram of the stand drive is shown in Fig. 1. Analysis of the rolling torque for each stand of the rolling mill (Table 1) shows that, in the first and subsequent stands, the peak torques decline on account of decrease in the strain and decrease in the tension on pipe capture by the rollers of the first stands. In steady reduction, the first (braking) stands operate in generator mode. In the intermediate stands, the peak torques and steady torques are distributed nonuniformly in some cases, on account of the nonuniform tension. In the last two or three stands, the steady torques increase because rear tension predominates [1]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000582_978-3-030-20131-9_325-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000582_978-3-030-20131-9_325-Figure1-1.png", + "caption": "Fig. 1. Parametrization of wheel and rail surfaces", + "texts": [ + ", any point of their surfaces can be defined by two coordinates [6]. Hence, the surface of each rail can be obtained through the sweep of its crosssection along a given path which is represented by a set of nodal points. These nodal points specify the position and orientation of the rail as function its arclength, r sides . The position of the origin of the rail profile is defined by r sider , and its orientation is given by a set of vectors r sidet , r siden and r sideb , which are the tangent, normal and binormal vectors, respectively, as can be depicted in Fig. 1. Then, the rail profile is represented by a two-dimensional function, in which r sideu is the surface parameter that defines the lateral rail coordinate and r sidef denotes the dependent vertical coordinate. In turn, since the wheel is a revolution body, the surface is defined by the rotation of its cross-section about its axis. Thus, any point in the wheel surface is characterized by the angular position w sides , and the lateral coordinate w sideu . In a similar manner, the wheel profile is represented by a two-dimensional function, in which w sidef denotes the dependent wheel profile coordinate. Note that the superscript side can be defined Based on this surface parametrization, Figure 1 demonstrates how two arbitrary points, P and Q, can be defined in the rail and wheel surfaces, respectively. In that sense, the position of point P can be expressed as r r, side side side P Pr r r where r sider denotes the position vector of the origin of the rail profile and depends on the rail arclength r sides , and r, side Pr is the distance vector between the rail profile origin and point P. In what concerns to the wheel, the position of an arbitrary point Q in its surface can be expressed, according Fig. 1, in the following manner ws w w, side side side Q Qr r r r (2) where wsr denotes the position vector of the wheelset center of mass, w sider represents the distance between the wheelset center of mass and the wheel profile origin, and w, side Qr is the distance vector between the wheel profile origin and point Q. The mathematical definition of the wheel and rail profiles assumes an important role on the study of wheel-rail contact interaction and can be performed in several different ways. As discussed in section 1, some authors neglect the transition zone between the tread and the flange in order to avoid a negative curvature zone, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure36.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure36.1-1.png", + "caption": "Fig. 36.1 Complete process of FDM process machine set-up", + "texts": [ + " Some researchers have also analysed the effect of 3D printing in Engineering education which is suitable for classrooms and public spaces and encourages action-based learning, exploration and innovation, to accelerate the acceptance and involvement of 3D printing technologies in the manufacturing sector [12, 13]. 416 P. Yadav et al. Not many people have focused the research on the effect of infill density on the mechanical properties of PLA materials for different raster orientations. Therefore, the present work focuses on the effect of infill density on mechanical properties in XY and XZ orientation and its fractography analysis using scanning electron microscopy (SEM). One of the most exciting and widely used additive manufacturing technologies is fused deposition modelling (Fig. 36.1) which is based on the material extrusion process. In FDM process, the polymer filament is fed to the liquefier chamber where it gets heated to a particular temperature depending upon the melting point of the material used for printing. Once the material is heated, it gets softens and melts and after that molten material is pushed through the nozzle from the liquefier with some suitable velocity. The extruded material in molten form is deposited on the platform layer by layer which slowly gets cooled at room temperature" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003455_1350650120954943-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003455_1350650120954943-Figure1-1.png", + "caption": "Figure 1. ABB IRB 6620 Configuration.38", + "texts": [ + " Articulated manipulators are prevailing and popular in industrial scenarios, and thus their motion control problem is of great significance. Take ABB IRB 6620, a typical six axes commercial industrial manipulator as an example for the case study, we can obtain insights about the effects of friction on dynamics and motion control, taking the complex environmental factors such as load and temperature into consideration. Kinematics of ABB IRB 6620 is first analyzed, and its modified D-H parameters are obtained from the CAD model provided at ABB\u2019s official website, which is shown in Figure 1. Figure 2 demonstrates the kinematics diagram of all six joints in ABB IRB 6620 manipulator adopting a modified D-H style, and its detailed kinematics parameters are displayed in Table 1, where the last column shows the initial position. With inertial parameters such as the center of mass and mass of links taken from the CAD model, we could obtain the dynamic model without friction first, by adopting iterative Newton-Euler dynamics algorithm2: s \u00bc M\u00f0H\u00de \u20acH \u00fe C\u00f0H; _H\u00de _H \u00fe G\u00f0H\u00de (1) where \u20acH; _H;H are joint acceleration, velocity and position respectively, s is driving torque applied by motor, M\u00f0H\u00de and C\u00f0H\u00de stand for inertial and centrifugal and Coriolis matrix respectively and G\u00f0H\u00de represents the gravitational torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001471_j.engfailanal.2019.104223-Figure14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001471_j.engfailanal.2019.104223-Figure14-1.png", + "caption": "Fig. 14. Metal bracket supporting upper centre cowl.", + "texts": [ + " Table 3 illustrates high strains at corresponding locations depicting the severity of both the events. High strain is observed during dynamic loading in shaker without pillion load. This shows that pillion load transfer shall not be the cause of failure. However, since the strain values are higher for pillion static load, it cannot be neglected and countermeasures should be taken. A metal stay is designed to support upper centre cowl in order to withstand pillion load in case of contact. Mountings of stay are between PTC bracket and upper centre cowl mounting with taillight as shown in Fig. 14. Strain measurement is acquired for static pillion load along with taillight stay. Strain has reduced, however presence of strain depicts that the added stay does not stop pillion load to get transferred to rear cowl assembly. Thus, impression test is performed again and same is confirmed. Therefore, to avoid contact between seat base & upper centre cowl, different counter-measure is proposed in which a cross metal pipe as shown in Fig. 15 is mounted over rear grip. It provides additional support to seat base and constrains its downward deflection" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure12-1.png", + "caption": "Figure 12 Sixth mode", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0000724_978-981-13-3305-7_190-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000724_978-981-13-3305-7_190-Figure4-1.png", + "caption": "Fig. 4. The continuous transition strategy", + "texts": [ + " Most of these disturbances and uncertainties can be reduced and even eliminated through the method of robust control, except the large AOA aerodynamic effect, which happens in the procedure of the discontinuous transition mode. Fortunately, it can be reduced by changing the discontinuous strategy into continuous transition strategy shown in Fig. 3, which means that the magnitude of AOA is kept roughly the same during the whole transition procedure, meanwhile, the flight path is a continuous curve in the longitudinal plane. However, compared with discontinuous transition strategy, continuous transition strategy shown in Fig. 4 requires higher level of state estimation accuracy and flight performance to maintain the AOA at a constant value. Besides, more effective and powerful propulsion system is needed to obtain the larger tangential acceleration of the flight path to reach the minimum flight speed of fixed-wing mode in a relatively short time. In this paper, the discontinuous transition strategy is used because of its easy realizability and simple flight parameter adjustment procedure, and the continuous transition strategy will be determined in the future" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001911_ijmmme.2020040104-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001911_ijmmme.2020040104-Figure5-1.png", + "caption": "Figure 5. Deformation of each model", + "texts": [], + "surrounding_texts": [ + "The\ufeffpoint\ufeffA\ufeffin\ufeffFigure\ufeff5\ufeffrefers\ufeffto\ufeffthe\ufeffposition\ufeffvector\ufeffof\ufeffa\ufefftip\ufeffof\ufeffthe\ufeffsoft\ufeffactuator\ufeffbefore\ufeffpressurized\ufeff and\ufeffpoint\ufeffB\ufeffrefers\ufeffto\ufeffthe\ufefffinal\ufeffposition\ufeffvector\ufeffof\ufeffthe\ufeffsame\ufeffpoint\ufeffafter\ufeffthe\ufeffactuation.\ufeffThe\ufeffstrain\ufeff(%)\ufeffof\ufeff the\ufeffcontraction\ufeffand\ufeffextension\ufeffcan\ufeffbe\ufeffcomputed\ufeffas: \u03b5 = \u00d7 Change in length Original length 100 \ufeff The\ufeffbending\ufeffangle\ufeffof\ufeffthe\ufeffbending\ufeffactuator\ufeffcan\ufeffbe\ufeffcomputed\ufeffas: \u03b8\ufeff=\ufeff tan tan \u2212 \u2212 \u2212 < + \u2212 > 1 1 231 231 90 231 23 x y if y y x if y \u2206 \u2206 1 \ufeff Figure\ufeff6\ufeffshows\ufeffthe\ufeffstrain\ufeffin\ufeffcontraction\ufeffactuator\ufeffand\ufeffextension\ufeffactuator\ufeffof\ufeffboth\ufeffmaterials.\ufeffAs\ufeff expected,\ufeffthe\ufeffstrain\ufeffof\ufeffboth\ufeffextension\ufeffand\ufeffcontraction\ufeffactuators\ufefflooks\ufeffto\ufeffbe\ufeffnonlinear.\ufeffIn\ufeffregard\ufeffto\ufeff actuators\ufeffof\ufeffP1\ufeffSilicone\ufeffRTV,\ufeffthe\ufeffsudden\ufeffincrease\ufeffin\ufeffthe\ufeffstrain\ufeffis\ufeffobserved,\ufeffas\ufeffits\ufeffstiffness\ufeffis\ufefflower\ufeff than\ufeffSilicone/MWCNT.\ufeffThe\ufeffstrain\ufeffin\ufeffboth\ufeffcontraction\ufeffactuator\ufeffand\ufeffextension\ufeffactuator\ufeffis\ufefflinear\ufeffup\ufeff to\ufeffabout\ufeff50\ufeffkPa,\ufeffof\ufeffthe\ufeffinput\ufeffpressure.\ufeffThe\ufeffmaximum\ufeffstrain\ufeffin\ufeffextension\ufeffactuator\ufeffmade\ufeffof\ufeffSilicone\ufeff RTV\ufeffis\ufeff70%,\ufeffwhich\ufeffis\ufeffobserved\ufeffto\ufeffbe\ufeffthe\ufeffmaximum\ufeffamong\ufeffall\ufeffother\ufeffactuators.\ufeffThe\ufeffmaximum\ufeffstrain\ufeff in\ufeffcontraction\ufeffactuator\ufeffmade\ufeffof\ufeffSilicone\ufeffRTV\ufeffis\ufeff30%.\ufeffIt\ufeffis\ufeffalso\ufeffobserved\ufeffthat\ufeffit\ufefftakes\ufeffthe\ufeffmaximum\ufeff of\ufeff30%\ufeffat\ufeffabout\ufeff100\ufeffkPa\ufeffand\ufeffthen\ufeffdescends\ufeffslightly\ufefflow.\ufeffIt\ufeffis\ufeffan\ufeffinteresting\ufeffphenomenon\ufeffnoticed\ufeffin\ufeff the\ufeffsimulation\ufeffas\ufeffthere\ufeffis\ufeffa\ufefflow\ufeffchange\ufeffin\ufeffcontraction\ufeffbeyond\ufeff100\ufeffkPa.\ufeffThe\ufeffmaximum\ufeffinput\ufeffpressure\ufeff the\ufeffextension\ufeffactuator\ufeffand\ufeffcontraction\ufeffactuator\ufeffcan\ufeffwithstand\ufeffare\ufeff520\ufeffkPa\ufeffand\ufeff200\ufeffkPa\ufeffrespectively. In\ufeff regard\ufeff to\ufeff extension\ufeff actuator\ufeff and\ufeff contraction\ufeff actuator\ufeff made\ufeff of\ufeff Silicone/MWCNT,\ufeff the\ufeff strain\ufeff is\ufeff observed\ufeff increasing\ufeff gradually\ufeff upon\ufeff increase\ufeff in\ufeff the\ufeff input\ufeff pressure.\ufeff The\ufeff maximum\ufeff strain\ufeffin\ufeffextension\ufeffactuator\ufeffis\ufeffnoticed\ufeffto\ufeffbe\ufeff40%\ufeffand\ufeffthe\ufeffstrain\ufeffin\ufeffcontraction\ufeffactuator\ufeffis\ufeff30%.\ufeff Both\ufeffof\ufeffthese\ufeffactuators\ufeffare\ufeffable\ufeffto\ufeffwithstand\ufeff1000\ufeffkPa.\ufeffIt\ufeffis\ufeffnot\ufeffable\ufeffto\ufefffind\ufeffthe\ufeffmaximum\ufeffinput\ufeff pressure\ufeffdue\ufeffto\ufeffthe\ufefflimitation\ufeffof\ufeffthe\ufeffhardware\ufeffand\ufeffsoftware\ufeffin\ufeffthe\ufefflaboratory.\ufeffFrom\ufeffthese\ufeffresults,\ufeff it\ufeffis\ufeffconcluded\ufeffthat\ufeffSilicone\ufeffRTV\ufeffis\ufeffbetter\ufeffin\ufeffgetting\ufeffthe\ufeffhigh\ufeffstrain\ufeffand\ufefffast\ufeffresponse.\ufeffSilicone/ MWCNT\ufeffis\ufeffbetter\ufeffin\ufeffperforming\ufeffa\ufefflong\ufeffactuation. Figure\ufeff7\ufeffshows\ufeffthe\ufefftrajectory\ufeffof\ufeffthe\ufeffbending\ufeffactuator.\ufeffThe\ufeffbending\ufeffactuator\ufeffis\ufeffdesigned\ufeffwith\ufefftwo\ufeff angles;\ufeffone\ufeffrelated\ufeffto\ufeffcontraction\ufeffand\ufeffother\ufeffone\ufeffrelated\ufeffto\ufeffextension.\ufeffWhen\ufeffthe\ufeffrubber\ufefftube\ufeffis\ufeffrestricted\ufeff to\ufeffhave\ufeffextension\ufeffand\ufeffcontraction\ufeffat\ufeffa\ufefftime,\ufeffthe\ufeffdisplacement\ufeffbecomes\ufeffbending.\ufeffThe\ufeffdisplacement\ufeff in\ufeffY-axis\ufeffis\ufeffthe\ufeffextension\ufeffof\ufeffthe\ufefftube\ufeffwhereas\ufeffthe\ufeffdisplacement\ufeffin\ufeffZ-axis\ufeffis\ufeffthe\ufeffcontraction\ufeffof\ufeffthe\ufeff tube.\ufeffDue\ufeff to\ufeff the\ufeffextension\ufeffof\ufeff the\ufeff tube\ufeffand\ufeffcontraction\ufeffoccur\ufeffat\ufeffa\ufeff time,\ufeff the\ufeffactuation\ufeff takes\ufeffplace\ufeff in\ufeffthe\ufeffelliptical\ufeffpath,\ufeffwhich\ufeffregarded\ufeffas\ufeffbending\ufeffof\ufeffthe\ufeffactuator.\ufeffFigure\ufeff8\ufeffand\ufeffFigure\ufeff9\ufeffshow\ufeffthe\ufeff bending\ufeffangle\ufeffand\ufeffthe\ufeffcurling\ufeffeffect\ufeffof\ufeffsilicone\ufeffRTV\ufeffand\ufeffsilicone/MWCNT\ufeffactuators.\ufeffAs\ufeffexpected,\ufeff the\ufeffactuator\ufeffof\ufeffSilicone\ufeffRTV\ufeffshows\ufeffthe\ufefffast-sharp\ufeffincrease\ufeffin\ufeffbending,\ufeffreaches\ufeffmaximum\ufeffof\ufeff84\u00b0\ufeffand\ufeff then\ufeffstarts\ufeffsharp\ufeffdecrease.\ufeffThe\ufeffdecrease\ufeffin\ufeffbending\ufeffangle\ufeffis\ufeffdue\ufeffto\ufeffcurl\ufeffitself\ufeffinwards.\ufeffOn\ufeffthe\ufeffother\ufeff hand,\ufeffSilicone/MWCNT\ufeffactuator\ufeffshows\ufeffthe\ufeffgradual\ufeffbending\ufeffup\ufeffto\ufeffthe\ufeffmaximum\ufeffbending\ufeffangle\ufeffof\ufeff 128\u00b0\ufeffat\ufeff550\ufeffkPa\ufeffand\ufeffthen\ufeffstarts\ufeffdecrease.\ufeffBut\ufeffthe\ufeffinteresting\ufeffphenomenon\ufeffnoticed\ufefffrom\ufeffit\ufeffis\ufeffthat\ufeffthere\ufeff is\ufeffnot\ufeffmuch\ufeffbending\ufeffbeyond\ufeff580\ufeffkPa,\ufeffwhich\ufeffconsidered\ufeffto\ufeffbe\ufeffno\ufeffsignificant\ufeffimpact\ufeffof\ufeffcurling\ufeffeffect." + ] + }, + { + "image_filename": "designv11_80_0001471_j.engfailanal.2019.104223-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001471_j.engfailanal.2019.104223-Figure7-1.png", + "caption": "Fig. 7. Strain gauge applied to measure assembly strain.", + "texts": [ + " (a) Frame mounting location (b) Rear lower centre cowl supported with rear fender. noticed that there is no stress at failure location as shown in Fig. 6. Thus, deviation in frame mountings is not a cause of failure. However, the deviations are communicated to frame manufacturing plant and it is controlled by improving the design of fixtures. Rear cowl, rear lower centre cowl, upper centre cowl and tail light are part of a sub-assembly as shown in Fig. 1. Strain gauge are pasted in locations shown in Fig. 7 to capture assembly stresses. Ideally, assembly strains are expected to be close to zero, if the parts are as per design. However, high assembly strain is observed near mounting location of rear lower centre cowl during sub-assembly as well as vehicle assembly which could result in stress concentration. Table 2 illustrates maximum strain at corresponding locations. High assembly strain developed could be because of warpage in the moulded plastic part. Part warpage occurs when there is more shrinkage in the thickness direction (corner effect), and/or non-uniform shrinkage from side to side due to temperature difference through thickness (differential cooling), and/or higher shrinkage due to varying thickness, as thick zones take longer to cool (differential shrinkage)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003377_ccdc49329.2020.9164095-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003377_ccdc49329.2020.9164095-Figure1-1.png", + "caption": "Figure 1: The Folding Wing-Tip UAV Configuration", + "texts": [ + " The extra force Fext and moment Mext originated from morphing are expressed as follows: Fext=\u2212md\u0308cm\u2212m[ \u00af\u0307\u03c9]dcm\u2212m[\u03c9\u0304]2dcm\u2212 2m[\u03c9\u0304]d\u0307cm, (2) Mext =\u2212 J\u0307\u03c9 \u2212m[ \u00afdcm](\u03c9\u0304Vb + V\u0307b) \u2212 \u222b [d\u0304]d\u0308dm\u2212 [\u03c9\u0304] \u222b [d\u0304]d\u0307dm, (3) where the term dcm denotes the position of the CM within body frame and d is the position of a mass element. The term [d\u0304] and [ \u00afdcm] represent the 3-dimension skew-symmetric matrix representation of d and dcm, respectively. In the process of morphing, various dynamic properties change in the body frame such as lift coefficient, drag coefficient and moment coefficient. The centroid dynamics equations above can vary in different aircrafts, so figuring out the wing-shape parameters of folding wing-tip UAV is essential. The top view and the front view of the folding wing-tip UAV can be given in Figure 1: As is shown in Figure 1, the aircraft consists of a fixed fuselage, two fixed inner wings and two foldable symmetrical wing tips. The dihedral angle \u03b41 and \u03b42 represent the folding angle of wing tips. Similar to the simplifications proposed in [19], some assumptions are proposed to compute the additional force and moment caused by folding of wing tips: 1) Two foldable symmetrical wing tips have the same span ltip and the same mass mtip, and two fixed inner wing has the same span lf and the same mass mf ; 2) The folding angular rate during morphing keeps constant; 3) The folding of wing tip can be done in a short time, and the start time tj and the finish time tk of the folding process have the relationship tj \u2248 tk" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001549_s11223-019-00110-8-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001549_s11223-019-00110-8-Figure9-1.png", + "caption": "Fig. 9. Thermal strain diagram of the ring gear at: (a) 50 ; (b) 100 .", + "texts": [], + "surrounding_texts": [ + "3.1. Design of the Thermoelastic Analysis Model of the Meshing Pair. During the meshing process of the movable tooth and ring gear, the temperature produced is higher than that of the wave generator and movable tooth due to different friction modes. To investigate the variation law of the maximum thermal strain of the meshing pair, the influence of the heat source intensity caused by the sliding friction between the ring gear and the movable tooth on the thermal strain and stress can be analyzed. During the meshing process, the interaction and the relative sliding speed between the ring gear and movable tooth produce frictional heat which is assigned to the ring gear and movable tooth according to a certain proportional distribution coefficient. The meshing pair temperature raises and the thermal strain takes place. A contact half-width between the ring gear and movable tooth is less than their thickness. The friction contact area between the ring gear and movable tooth is rectangular. The solid material is assumed to be isotropic and the internal heat transfer is uniform. The task can be solved via a 2D model. Hot deformation of the meshing pair can be calculated by Green\u2019s function method. A heat source in the rectangular area can be transformed into a linear heat source in the cross-section direction. A new general solution of Green\u2019s function on a semi-infinite plane surface can be applied to a linear heat source. As shown in Fig. 6, the ring gear middle plane is regarded as a two-dimensional semi-infinite plane, where the normal line is the heat source depth direction. In the two-dimensional Descartes coordinate system, constitutive equations of the isotropic thermal elastic material have the following form: x x y T y y x T G u x u y T G u y u x ( ) , ( ) 2 2 T G u x u y xy x y , , (11) where ux and u y are displacement components in the x- and y -directions, respectively, x , y , and xy are normal stress and shear stress components, T is the temperature increment, T is the thermal modulus, and and G are the Lame constant and the shear modulus, respectively. The total thermal strain of a point in the semi-infinite plane is u u ux y 2 2 . (12) The equivalent thermal stress is 2 2 6 2 2 ( ) .x y xy (13) Using the meshing parameters in Table 1 and Eqs. (12) and (13), the thermal strains and thermal stresses at the angle of 100 were computed via MATLAB. As shown in Fig. 7, the heat deformation near the heat source is the smallest. The maximum thermal strain is 7.12 m at the location far away from the heat source. As can be seen from Fig. 8, the thermal stress near the heat source reaches a maximum of 114 MPa. It is shown that the location of the heat source has a great influence on the distribution of thermal stress. The thermal strain distribution depends not only on the heat source location but also on the tooth profile and material properties. 3.2. Thermal Structure Coupling Analysis Based on ANSYS Workbench. We selected the heat flux values as 176 and 438 kW/m 2 , and the corresponding rotation-angle of the wave generator 50 and 100 . Thermal strain and stress diagrams obtained by simulation are shown in Figs. 9 and 10. The maximum thermal strain always appears near the top of the ring gear tooth and the minimum appears near the root of the tooth. This is because thermal expansion of a part relates to its shape. The radius of curvature of the tooth top area is negative, and the radius of curvature of the tooth root is positive. The thermal strain reduces the gap between the movable tooth and the ring gear and the movable tooth may be jammed. As can be seen from Fig. 10, the maximum thermal stress is always near the root of the tooth, while the minimum is near the tooth\u2019s top. This is because the radius of curvature of the tooth root is negative, so stress concentrates at this place. Therefore, the tooth root is a dangerous place of the thermal stress damage during the meshing of the movable tooth and the ring gear at high speed and heavy load conditions." + ] + }, + { + "image_filename": "designv11_80_0001227_rusautocon.2019.8867687-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001227_rusautocon.2019.8867687-Figure1-1.png", + "caption": "Fig. 1. A simplified scheme of laboratory installation", + "texts": [ + " The measurement of the shaft angular speed is performed using an incremental encoder OsiSense XCC. The main crane characteristics are specified in Table 1. The trolley reference speed has a trapezoidal shape. During the experiment, a video of the crane was recorded with a simultaneous oscilloscope procedure of the electric drive variables. The speed oscillogram is obtained with the help of SoMove software for parameterizing the frequency converter. The scheme of the laboratory installation for video recording is shown in Fig. 1. The tracking of the bridge crane trolley is performed in the Cartesian system of coordinates. The tracking of the trolley movement begins at the initial position XOY, the trolley moves along the x and y-axes. The camera is installed exactly in the middle of the crane at the height on the level of the crane trolley. The distance from the beginning of the crane railways to the installation site of the camera is selected so that the beginning of the crane railways lay into the area of view of the camera" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.15-1.png", + "caption": "Figure 1.15. Parametrization for the direct geometric model of a manipulator robot", + "texts": [ + " Verify that the drawing is correct by moving the six degrees of freedom of the robot one by one. Obtain a three-dimensional representation such as the one illustrated in Figure 1.14. 3) By using the relation $|R0 = $ % cos ! cos' \" sin' 0 cos ! sin' cos' 0 \" sin ! 0 1 & ' $ % &\u0307 !\u0307 '\u0307 & ' , draw the instantaneous rotation vector of the robot. 4) Simulate the robot in various conditions using a Euler method. EXERCISE 1.11.\u2013 Manipulator robot A manipulator robot, such as Staubli (represented on Figure 1.15), is composed of several rigid arms. We retrieve the coordinates of the end effector, at the extremity of the robot, using a series of geometric transformations. We can show that a parametrization with four degrees of freedom allows us to represent these transformations. There are several possible parametrizations, each with its own advantages and disadvantages. The most widely used one is probably the Denavit\u2013Hartenberg parametrization. In the case where the articulations are rotational joints (as is the case of the Staubli robot where the joins can turn), the parametrization represented by the figure might prove to be practical since it makes it easier to draw the robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002364_022108-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002364_022108-Figure4-1.png", + "caption": "Fig. 4 6 DOF gear dynamic model", + "texts": [ + "1088/1755-1315/440/2/022108 The results indicate that by introducing tooth root crack, mesh stiffness of the helical gear pair decreases significantly. And the reduction of the mesh stiffness increases with the growth of the crack depth. Based on the mesh stiffness of helical gear pair with tooth crack, a 6 DOF gear dynamic model is established to investigate the vibration characteristics of helical gear transmission system with and without tooth root crack. In this paper, the effect of the friction force is neglected. The dynamic model is shown in Fig. 4 and the equations are displayed as follows. p tan tan p p pb gg g gb p pp b pBx p pBx g gg b gBx g gBx p pBy p pByp p g gBy g gByg g J T N R J T N R m x N K x C x m x N K x C x m y N K y C y m y N K y C y (16) ESMA 2019 IOP Conf. Series: Earth and Environmental Science 440 (2020) 022108 IOP Publishing doi:10.1088/1755-1315/440/2/022108 Where p J / g J is the inertial moment of pinion/gear; p m / g m is the mass of pinion/gear; p T / g T is the external torque; pBxK / gBxK , pByK / gByK , pBxC / gBxC , pByC / gByC is the stiffness and damping in x and y direction of the supporting bearings" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure5-1.png", + "caption": "Fig. 5. One step sloting model", + "texts": [], + "surrounding_texts": [ + "Using the FEMM 4.2 and coupled with LUA 4.0 scripting, the PMMs characteristics were investigated [7],[11],[12]-[14]. By the combination of FEMM and LUA 4.0 could increase a quick execution for the implementation of a complete simulation of a specific PMMs. Another advantage of LUA application is the capability of parallel computation can be achieved. At the beginning of each simulation, the simulated of PMMs structure were generated in Auto-CAD then exported to the FEMM file. Comparisons of air gap magnetic flux distribution and CT for the PMMs studied were investigated. This means that the proposed PMM model (twosteps slot) was promising for the presence of two steps of slotting in the magnets that do not distort the balance of magnetic force in the air gap of the machine of one-step model. Figures 4, 5 and 6, shows that the value of the flux distribution due to the changing in the magnet structure will not destroy the value of the machine's core losses, as it could be observed that all the value of the flux density approximately 1.436 Tesla. It has been found that if it implemented to the proposed method, it will not change the core losses of the proposed model." + ] + }, + { + "image_filename": "designv11_80_0001442_042023-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001442_042023-Figure3-1.png", + "caption": "Fig 3. The schematic of five-axis linkage", + "texts": [ + " Through the transition of the guiding plate, the insulated hook can be hung on the line, and the inside of the insulated hook head is inlaid with a small-sized bearing. By pulling implemented by operators on the ground, the device can be moved smoothly to the foreign objects. The transmission part adopts five-axis linkage, including two power shafts and three transmission shafts. The outer surface of the power shaft is polished to increase the external friction. The insulated rope is deployed according to the fig 3. The operator drives the blade on the power shaft to rotate by pulling the insulating rope. The cutting blade adopts the design of the drum blade. The longitudinal length of the blade is 120mm. The blade mounting hole distance can meet the requirements of ordinary utility knife size and is easy to be replaced. The tool cross-section rotation radius is 55mm, which can avoid the disadvantages of the flat blade design, which is easy to be entangled. ICAMMT 2019 IOP Conf. Series: Materials Science and Engineering 631 (2019) 042023 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure8-1.png", + "caption": "Fig. 8 Chiral structure CAD.", + "texts": [ + " The auxetic effect is observed as the diagonal ribs aligned along the horizontal direction move apart in the vertical direction under tension. Tests have shown that most of the structures involving re-entrant honeycombs undergo deformation. A chiral formation is defined as a nonsuperimposable mirror image. Such structures exhibit a Poisson\u2019s ratio of close to 1 (Saxena et al., 2016).When the ligaments are pulled (i.e., a longitudinal strain is applied), the rigid ring rotates, causing unwinding of the ligaments, which leads to its auxetic behavior. As shown in Fig. 8, we designed and fabricated a tetra-chiral structure with each central node connected to four other nodes. While we wanted to construct a hexachiral structure, the geometric constraints set by our design specifications did not allow us to construct such designs as the tube diameter would have had to be increased to 20mm. A star honeycomb structure is a subset of re-entrant structures and a variant of the traditional hexagonal re-entrant design. As is common with re-entrancy, the deformation of the star structure is through the hinging of the walls of the cells, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure78.9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure78.9-1.png", + "caption": "Fig. 78.9 Equivalent (von Mises) stresses of a existing and b modified FH", + "texts": [], + "surrounding_texts": [ + "The test is carried out on each sample at five different locations with three reputations. These results show that the hardness of the material is not the same throughout the hammer. At the hard-facing welded part, the hardness is high and less at the tail part. Due to this, the Fibrizer hammer frequently breaks at the welded part. The welding heat input is important welding parameter, which effects on the structure and properties of the weld metal. At location-1, sample harness result shows that the hardness varied between 110 and 124 VHN. At the location-2, Fibrizer hammer hardness is between 146 and 151 VHN, and at the location-3, hardness is between 152 and 167 VHN. 78.3.4 Cost Comparison Between Existing and Modified Fibrizer Table 78.6 indicates the cost comparison between existing and modified Fibrizer. The difference in cost between them is $11,088.00. 78.3.5 Harmonic Analysis Results of Existing and Modified Fibrizer Hammers ANSYS finite element software is used for the simulation to performing harmonic analysis. From the simulation results, total deformation, direction deformation, von Mises stresses, maximum shear stresses, and maximum amplitude are evaluated for existing and modified Fibrizer hammers. 942 T. Mathewos et al. From Figs. 78.7, 78.8, 78.9, and 78.10 analysis, results are tabulated in Table 78.7. From the results, it is understood that the total deformation in the modified Fibrizer hammer is lesser than the total deformation of the existing hammer. This shows that the modified Fibrizer hammer is more reliable than the current one. As the result shows that the maximum equivalent (von Mises) stress of current FH is increased by double, the higher the stress, the more the material will be exposed to be broken. The maximum shear stress result in the current Fibrizer hammer is more than the modified Fibrizer hammer. The smaller the radius (r) and web (h) in stepped plate of the Fibrizer hammer, the higher will be the stress. In Table 78.7, analysis results show that there is less total and directional deformation in modified Fibrizer hammer. The equivalent (von Mises) stress and max. shear stress result also much less by half from the existing. Similar manner frequency responses are verified for other surfaces also and tabulated in Table 78.8. From Table 78.8 data, the output of harmonic response has been taken in the form of amplitude which can be understood as mean stress development in any engineering components. Modified FH has more amplitude initially and gradually decreased compared to current FH. The overall behavior modified FH is taking less deformations and stresses by which life span of FH definitely will improve. 78 Design Analysis and Modification of Sugarcane Fibrizer Hammer \u2026 943 944 T. Mathewos et al." + ] + }, + { + "image_filename": "designv11_80_0001739_ismsit.2019.8932756-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001739_ismsit.2019.8932756-Figure4-1.png", + "caption": "Fig. 4. Y-Axis", + "texts": [], + "surrounding_texts": [ + "978-1-7281-3789-6/19/$31.00 \u00a92019 IEEE\nKeywords\u2014 automation, gardening, mobile application, timely monitoring\nI. INTRODUCTION\nIn manufacturing sector, to control machine tool a processed is used which is called CNC machining. The process includes the use of computers to control the machine tools. Tools can be any object which helps us in performing different task like extruder in 3D printer, blade in lathe machine, drill bit in milling machine. CNC stands for Computer Numerical Control. It look like an ordinary personal computer (PC), but it differs due to unique software and console to control the machine.\nTo control the speed, location, coordinates and feed rate of machine a specialized CNC machining language is used, which is known as G-code. With the help of CNC machining we can control the position and velocity with great precision. CNC machining is currently being used in manufacturing for plastic and metal parts.\nThere are lots of advantages of using CNC machines. It\u2019s hard to manufacture any part in manual machining. Manual machining consumes more time and energy and it does not have much precision. While using CNC machining a lot of man power and time is secured and also the parts manufactured with CNC machining has great precision.\nA CNC machine consist multiple individual motors, which helps its tool to move in respective specific direction. A CNC Machine consist of two or more axis of motion. Usually CNC machine which are being used are consist 3-axis of motion. If there is any additional rotational movement then it will be 4-axis.\nII. LITERATURE REVIEW\nCommercially, Agriculture has been reached a newly high level of automation up till now, mainly for growing\ncrops on broad land. Fine-grain satellite imagery is also available commercially for pesticides and fertilizer related applications, leading to the precision agriculture\u2019s novel paradigm, its basic goal is to save water and pesticides [1]. The interest is growing day by day in autonomous farming so we can move towards smaller robotic platforms that can work on individual basis and can done some manipulation in the field with the help of precise sensing [2].\nWith the increasing research in the field of computer vision [3] and mechatronics which is helping us to led to an autonomous solution for harvesting some specialty crops which includes Cherries [4], apples, tomatoes, cucumber, mushrooms, strawberries, melons and much other.\nAnother active research on which work is being done is automatic weed control. Grey-level vision is used to navigate in structured outdoor environment and uses color vision to differentiate between the required crops and weed which is needed to be removed. Beside their multiple applications in the field, we envision automated agriculture robot which is precise and able to work without any operator in any sort of environment like urban areas, house roofs or can easily work in harsh environment like outer space.\nAlso, a lot of work is done in this field like autonomous targeted spraying [5] to removes pests from crops which can destroy and can have an effect on crops production, design of optimum manipulator for autonomous de-leafing [6] process of cucumber plant to prevent it from fungal disease which is considered by its growers but is costly. In addition to all of this there are also some virtual experimentation frameworks which have been developed for agriculture robots, [7], [8] ForboMind is a customized software platform which was introduced to support and help field robot done agriculture task done with precision and to promote to reuse the robotics components.\nOn the other hand, Agriculture field robots contributes to improve soil health, yield and reliability of operation. Which are commonly equipped with multiple sensors and cameras for navigation, localization, mapping and path planning algorithm.\nAlso, farming industry make use of drones for surveillance and monitoring fields. These drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides.", + "III. SOCIO-ECONOMIC SIGNIFICANCE\nAt present farming industry make use of drones for surveillance and monitoring fields these drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides. Similar Machinery exists but is only household specific; however we are aiming to target research and development departments in pharmaceutical and food industry, where setting up small greenhouse is required. Further we would be using some tool head for multiple operations instead of relying on magnetism for tool selection. Cutting down cost is another objective by using alternative materials to steel.\nIV. DESIGN\nWhile developing the project we went through some major and minor issues which lead us to make multiple iterations in the design phase. In this section we will discuss these iterations.\n Iteration 1:\nThis is the initial idea and concept on which we started to work. Our initial idea was to have x-axis, y-axis, z-axis and one rotary axis for the axis of motion. We also thought about adding multi-head mounter and add moisture sensor, shower and seeder. Initially we want to use rack and pinion mechanism. Issues: Rack and pinion mechanism can increase the vibrations which can affect the structure.\nWe then start to work on to select a mechanism for x-axis. We start to work on the stepper motor and lead screw mechanism. Issues: Our project has an open base so lead screw can\u2019t be mounted at the bottom like every CNC milling machine and usage of two stepper motor for just one axis can effect or destroy the structure if something happens to one motor.\n Iteration 3: So we started to think more about it and started to work on another solution to move the x-axis with just one motor. We finally decide to move the whole axis with one motor by using belt and pulley mechanism as shown in figure below.\nIn this project we use t-slotted extruded bars for developing our structure. Y-axis consist of simple lead screw mechanism.\nZ-axis consist of simple lead screw mechanism for linear motion. On z-axis a servo motor is mounted for giving a rotary motion to the multi head. The final design after the iterations is shown below.", + " Vacuum Pump: When a signal is sent out from controller to the vacuum pump and vacuum pump turn on and suck air and move towards seed when it come near the seed, then seed float and got stuck on the nozzle until the vacuum is on. When it reaches to the desired location then the controller sends another signal and vacuum pump turns off and the seed drops on the location the grid.\n Water Pump: When a signal is sent out from the controller to the water pump then water pump turn on and sprinkle the water over the grid with the help of shower for some specific time and then another signal comes from the controller and turns it off. Piping and instrumentation diagram (P&ID) is shown below:\nCurrently there are lots of global challenges which we are facing like global warming, food needs, poverty, degradation of climate and much more. Sustainable development goals introduced by United Nation helps us address these problems and base on these problems find solution to make future better not just for humans but also for the creature that exist on globe. The goals which can be achieved by our project \u201cGarden Tech\u201d are as follows:\n Zero Hunger: As we all know that human population is increasing rapidly to accommodate them deforestation is taking place. It is effecting our climate and results in global warming due to which glaciers are melting. Flooding is taking place more often than ever before and leads to devastation of arable land. Due to these circumstances food growth is not increasing rapidly like human population. The anticipated population according to United Nations Department of Economic and Social Affairs (UN DESA) is shown below:" + ] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure9-1.png", + "caption": "Figure 9 Third mode", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0003328_012034-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003328_012034-Figure4-1.png", + "caption": "Figure 4. Machine modeled in SolidWorks.", + "texts": [ + " In the design of the pieces that form part of the mechanism and the modelling of the existing ones, the optimization of the materials that carry the quality line was considered, in addition it was stipulated as a biaxial system, emphasizing the automation of the welding speed. Consequently, the system is 6th International Week of Science, Technology and Innovation (6th IWSTI) Journal of Physics: Conference Series 1587 (2020) 012034 IOP Publishing doi:10.1088/1742-6596/1587/1/012034 composed horizontally of a base containing a power screw coupled to a servomotor and a pair of sliders with their respective blocks, supported on a table; vertically, as shown in Figure 4, it consists of a worm adapted to the base that holds the welding gun and two additional sliding shafts that provide manually adjustable freedom of movement. The study focused on the calculation of the main stresses of the power screw since it is the piece most exposed to alterations in its movement with loads. The equations used were deduced through the physical principles of the second law of newton, focusing on the forces present and the kinematics of torque. In this way, the results of the stress analysis corroborate the hypothesis that the forces involved in the movement of the power screw will not alter the speed of the process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002982_j.mechmachtheory.2020.104001-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002982_j.mechmachtheory.2020.104001-Figure4-1.png", + "caption": "Fig. 4. Two spherical concave contacts result in a revolute pair.", + "texts": [ + " In all the examples except in two, the contact geometries are identical sphere-to-sphere contacts. In the case of contact of locally spherical surfaces , k 1 m = k 2 m = k m and k 1 f = k 2 f = k f . Then, the above expressions for A and B in Eqs. (1) and (2) reduce to the following. A = \u22121 8 (k m + k f ) { L 2 + M 2 \u2212 (k m \u2212 k f ) LQ \u2217 + (k m \u2212 k f ) MP \u2217 \u2212 k m k f P \u22172 \u2212 k m k f Q \u22172 \u2212 (k m + k f ) R \u2217 d } (15) B = 1 4 (k m + k f ) 2 R \u2217 (16) 5.1. A pair with two contacts An object is in two point contacts with two fixed identical spheres as shown in Fig. 4 at contact- a and contact- b . As per the condition given in Fig. 3 (a), the points of contact are labeled C a and C b and are separated by a distance e along the common normal line. The shape of the object too at the contacts is locally spherical and identical at both the contacts. We show below that the situation corresponds to a revolute joint. k m and k f are the curvatures of the spheres at each contact of the movable and fixed object respectively. The contacts are concave with k m < 0 and k f > 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure3-1.png", + "caption": "Figure 3. Crank-rocker mechanism.", + "texts": [], + "surrounding_texts": [ + "The most common computational tool used for design and analysis of multi-body mechanical systems is ADAMS [45, 46]. Various types of links and joints along with their properties like mass, location of center of mass, moment of inertia and degrees of freedom can be assigned for a mechanism in this tool [18]. The joints defined by this tool are ideal as they do not have any clearance or deformation. Therefore, to compare performance of the mechanisms with joint clearances, clearances at joints were made in ADAMS. The clearance at revolute joint between coupler and slider has been made by making a cylindrical hole in the slider cube and attaching a cylinder to the end of the coupler. So, we can change the size of the hole to set a clearance size. Similarly, the clearance at revolute joint between crank and coupler has been made by attaching a cylinder to the end of the crank and making a hole in the attached cylinder. 3.1 Input factors Various input parameters for modeling of mechanism in ADAMS are as follows: 3.1a Clearance size: For a standard journal-bearing of journal diameter 20 mm, the clearance size ranges from 0.02 mm to 0.08 mm. However, due to wear during operation and other environmental factors, the clearance can increase. So, in this research work, the clearance size has been taken in the range of 0.02 mm to 1 mm. 3.1b Crank speed: To cover a wide range of cases, the speeds range from 100 rpm to 3000 rpm. 3.1c Contact conditions: For the modeling of contacts, ADAMS uses the contact method based on the impact function: IMPACT-Function-Based Contact. In this method, the solver computes the contact force from the IMPACT function available in the ADAMS function library. The normal force of the contact has two components: rigidity and viscous damping. The component of rigidity is a function of the penetration d. The component of the viscous damping is a function of the speed of penetration. In this model the normal force of contact is given as: FN \u00bc Kdn \u00fe STEP d; 0; 0; dmax;Cmax\u00f0 \u00de _d; d[ 0 0; d 0 \u00f07\u00de 3.1d Value of K: The revolute joint is a case of contact between two cylinders (one inside the other). So, the contact should start with a line contact and then become a 2D rectangular contact. But the value of K in this case will not only depend on the material and geometrical property but also the stress distribution between the cylinders, which cannot be determined accurately unless it is a static case and the force is applied externally. So, the researchers solved this problem by stating that the line contact in the revolute joint will only be present for two cylinders aligned with extreme precision. Also, a uniform force distribution over the length of the joint is not possible in real life conditions. Moreover, the forcedeformation diagrams for both spherical and cylindrical impact force models were studied in the literature [1\u20134] and it was found that the spherical and cylindrical forcedeformation diagrams are reasonably close. Based on these studies, we used the Hertzian contact force law between two spheres with the different parameters defined in Eq. (8). K \u00bc 4 3p hi \u00fe hj R1=2;R \u00bc RiRj Ri \u00fe Rj ; hk \u00bc 1 v2k pEk ; k \u00bc i; j \u00f08\u00de Ri, mi and Ei represent respectively the radii of the cylinders, the Poisson\u2019s ratio and the modulus of elasticity for element i. For clearance = 0.02 mm, journal radius = 10 mm and bearing radius = 10.02 mm E = 2.07*105 N/mm2; m = 0.29 Putting these values in the equation we get K = 3.37*105 N/m1.5 Similarly, for clearance = 0.1 mm, K = 3.377*105 N/m1.5 For clearance = 0.5 mm, K = 3.4*105 N/m1.5 3.1e Value of n:The value of n is usually taken to be 1.5 for metallic contacts. So, n = 1.5. 3.1f Value of damping coefficient: In ADAMS, the instantaneous damping coefficient is a cubic step function of the penetration given as: STEP d; 0; 0; dmax;Cmax\u00f0 \u00de \u00bc 0; d 0 Cmax d dmax 2 3 2 d dmax ; 0\\d\\dmax Cmax d dmax 8 >>< >>: \u00f09\u00de The value of Cmax should be approximately 1 percent of the value of K. dmax = 0.01 mm 3.2 Output factors Two factors were used as parameters for comparing the kinematic performance of different mechanisms, either displacement of the slider or angular rotation of the rocker attached to ground." + ] + }, + { + "image_filename": "designv11_80_0003693_icra40945.2020.9196516-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003693_icra40945.2020.9196516-Figure1-1.png", + "caption": "Fig. 1: Evaluation platform: Furuta pendulum.", + "texts": [ + " This objective function explicitly takes into account the waypoints and thus enables the generation of recognizable letter contours in long exposure photography. However, open-loop execution of the optimal trajectory is not sufficient on its own because even small deviations from the planned trajectory yield unrecognizable letters. Section IV details the implementation of the stabilizing feedback controller that enables efficient trajectory tracking. Finally, the resulting trajectories and letters are presented in Section V. The hardware platform used for experiments is the Quanser Qube shown in Fig. 1a. It implements the rotary inverted pendulum system introduced by Furuta et al. [8], which consists of a freely rotating pendulum attached to a motor-driven arm. A schematic is shown in Fig. 1b. While the arm can be rotated in the horizontal plane, the pendulum swings in the vertical plane orthogonal to the arm. The state of the nonlinear system is described by the two angles and the corresponding angular velocities x = [ \u03b8 \u03b1 \u03b8\u0307 \u03b1\u0307 ]T . The Furuta pendulum is a classical platform for evaluating control algorithms, appreciated for its rich passive dynamics and underactuation. Its equations of motion are provided in the Appendix, with derivations starting from the EulerLagrange equations available in [8] and [9]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001252_systol.2019.8864796-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001252_systol.2019.8864796-Figure2-1.png", + "caption": "Fig. 2: Two stages about appearance of disturbance torque \u03b8\u0307 \u2248 0) the rotational dynamic of hexacopter can be described as", + "texts": [ + " According to the rules, if one rotor is faulty, the symmetrical rotor should be stopped, because the torque generated by the symmetrical rotor can not be compensated effectively. The hexacopter hovers before the appearance of failure, the fault implies a stopped rotor. Assume that it occurs in the rotor r1. In the strategy of the standard degraded controller the rotor r4 is stopped after FDI [2]. For a better understanding of the generation of the disturbance torque \u03c4dis, it can be divided into two parts \u03c4dis,1 and \u03c4dis,2. It is necessary to pay attention to two stages as shown Fig. 2: \u2022 Stage 1: From the appearance of failure in rotor (1) to the detection and the isolation of failure (2), the time consumption is t1. The used controller in this stage is nominal controller, which is applied in faulty-free case. \u2022 Stage 2: After the detection and isolation of the fault (2), the rotor r4 is stopped. The degraded controller is applied to stabilize the position of the hexacopter (3), it takes the time t2 to complete the purpose. In nominal case the hexacopter hovers at a equilibrium point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000334_sii.2019.8700442-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000334_sii.2019.8700442-Figure7-1.png", + "caption": "Fig. 7: Brain model used in simulations which is reconstructed from MRI data of a porcine brain.", + "texts": [ + " 3) Failure When the damage coefficient d becomes bigger than a threshold dt , the element is considered to be fully fractured, and the element is removed to express fracture. The stiffness matrix is given by Ke(d) = 0 for d \u2265 dt . dt is set to 0.995. Saikali et al. have released MRI data of a porcine brain to the public [11], [12]. In this study, a geometric 3D model of the porcine brain was generated from the MRI data [12]. by using a software platform called 3D Slicer [13] The generated model is shown in Fig. 7 (a). The inside of the longitudinal cerebral fissure of the generated model is empty. The gap of the fissure was approximately 1.6 mm. Therefore, the inside of the fissure is filled with the model of connective tissues. A software platform called MeshLab [14] is used to fill the fissure by using the \u201cSurface Reconstruction: VCG\u201d function. Moreover, a slit to insert spatulas was made in the connective tissue. The brain model with connective tissue used in the simulation is shown in Fig. 7 (b). In the experiments presented in Section II, two spatulas were inserted into a slit made in the middle of the longitudinal cerebral fissure, and the Spatula 2 was moved while the Spatula 1 was fixed to split the fissure. To reproduce the experimental conditions, a similar slit was made on the connective tissue filled inside the cerebral fissure. In the simulations, the nodes on the left side of the slit were fixed and the nodes on the left side of the slit were set as the contact nodes to be moved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002568_0954407020909242-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002568_0954407020909242-Figure2-1.png", + "caption": "Figure 2. Force diagram of the vehicle plane structure.", + "texts": [], + "surrounding_texts": [ + "The proposed articulated engineering vehicle dynamics model is complex, and the subsystems are coupled and interacted with each other, significantly increasing the non-linearity of the entire system. The following assumptions are made for the model: (1) when the vehicle is running, the unevenness of the road surface remains constant, which is only related to the initial road grade; (2) keep driving at a constant speed during the course of driving; (3) ignore air, wheel roundness changes, and the driver\u2019s impact on the vehicle\u2019s stability; and (4) when the articulation angle is zero, the vehicle is bilaterally symmetrical. In this paper, a 16- DOF articulated engineering vehicle dynamic model considering the longitudinal vibration uxi, lateral vibration uyi, vertical vibration uzi, pitch angle ui, roll angle fi, yaw angle ci, and vertical vibration of four nonsuspended mass ui is established, as shown in Figures 1 and 2. A 16-DOF model of the wheel articulated loader is established using the Newton\u2013Euler method.14 The equations of motion are given below m1 \u20acux1 _uy1 _c1 =Fx1 +Fx2 +Cx \u00f01a\u00de m1 \u20acuy1 + _ux1 _c1 =Fy1 +Fy2 Cy \u00f01b\u00de m1\u20acuz1 = m1g Cz + Fz1 +Fz2\u00f0 \u00de + k1 u1 z1\u00f0 \u00de+ c1 _u1 _z1\u00f0 \u00de+ k1 u2 z2\u00f0 \u00de + c1 _u2 _z2\u00f0 \u00de \u00f01c\u00de Ix1\u20acf1 = Fy1 +Fy2 h1 + Fz1 Fz2\u00f0 \u00de0:5B1 + k1 u1 z1\u00f0 \u00de0:5B1 + c1 _u1 _z1\u00f0 \u00de0:5B1 k2 u2 z2\u00f0 \u00de0:5B1 c2 _u2 _z2\u00f0 \u00de0:5B1 +Cyhc1 My sinu \u00f01d\u00de Iy1\u20acu1 = Fx1 +Fx2\u00f0 \u00deh1 + Fz1 +Fz2\u00f0 \u00delf1 + k1 u1 z1\u00f0 \u00delf1 + c1 _u1 _z1\u00f0 \u00delf1 + k2 u2 z2\u00f0 \u00delf1 + c2 _u2 _z2\u00f0 \u00delf1 +Cxhc1 +Czlf2 My cosu \u00f01e\u00de Iz1\u20acc1 = Fx1 Fx2\u00f0 \u00de0:5B1 + Fy1 +Fy2 lf1 +Cylf2 \u00f01f\u00de The longitudinal, lateral, vertical, roll, pitch, and yaw equations of motion of the rear body are as follows m2 \u20acux2 _uy2 _c2 =Fx3 +Fx4 Cx cosu Cy sinu \u00f02a\u00de m2 \u20acuy2 + _ux2 _c2 =Fy3 +Fy4 Cx sinu+Cy cosu \u00f02b\u00de m2\u20acuz2 = m2g+Cz + Fz3 +Fz4\u00f0 \u00de+ k3 u3 z3\u00f0 \u00de + c3 _u3 _z3\u00f0 \u00de+ k4 u4 z4\u00f0 \u00de + c4 _u4 _z4\u00f0 \u00de \u00f02c\u00de Ix2\u20acf2 = Fy3 +Fy4 h2 + Fz3 Fz4\u00f0 \u00de0:5B2 + k3 u3 z3\u00f0 \u00deB2 + c3 _u3 _z3\u00f0 \u00de0:5B2 k4 u4 z4\u00f0 \u00de0:5B2 c4 _u4 _z4\u00f0 \u00de0:5B2 + Cy cosu Cx sinu hc2 \u00f02d\u00de Iy2\u20acu2 = Fx3 +Fx4\u00f0 \u00deh2 + Fz3 +Fz4\u00f0 \u00delr1 + k3 u3 z3\u00f0 \u00delr1 + c3 _u3 _z3\u00f0 \u00delr1 + k4 u4 z4\u00f0 \u00delr1 + c4 _u4 _z4\u00f0 \u00delr1 Cx cosu+Cy sinu hc2 Czlr2 +My \u00f02e\u00de Iz2\u20acc2 = Fx3 Fx4\u00f0 \u00de0:5B2 + Fy3 +Fy4 lr1 + Cy cosu Cx sinu lr2 \u00f02f\u00de Equations of motion of the four non-suspended masses in the vertical direction (front left, front right, rear left, and rear right, respectively) are as follows mf1\u20acu1 = kt1 q1 u1\u00f0 \u00de+ ct1 _c1 _u1\u00f0 \u00de k1 z1 u1\u00f0 \u00de c1 _z1 _u1\u00f0 \u00de \u00f03a\u00de mf2\u20acu2 = kt2 q2 u2\u00f0 \u00de+ ct2 _c2 _u2\u00f0 \u00de k2 z2 u2\u00f0 \u00de c2 _z2 _u2\u00f0 \u00de \u00f03b\u00de mf3\u20acu3 = kt3 q3 u3\u00f0 \u00de+ ct3 _c3 _u3\u00f0 \u00de k3 z3 u3\u00f0 \u00de c3 _z3 _u3\u00f0 \u00de \u00f03c\u00de mf4\u20acu4 = kt4 q4 u4\u00f0 \u00de+ ct4 _c4 _u4\u00f0 \u00de k4 z4 u4\u00f0 \u00de c4 _z4 _u4\u00f0 \u00de \u00f03d\u00de The generalized coordinates of the 16-DOF model are uT = ux1, uy1, uz1, ux2, uy2, uz2, u1, f,c1, u1,f2,c2, u1, u2, u3, u4 The dynamic equations are rewritten as follows M\u20acX t\u00f0 \u00de+C _X t\u00f0 \u00de+K t\u00f0 \u00deX t\u00f0 \u00de=Fz t\u00f0 \u00de+Fr t\u00f0 \u00de \u00f04\u00de where Fz and Fr represent the engine and road surface excitations, respectively. During the driving process, the external excitations mainly include road and engine excitations that are considered in our dynamic model. In this paper, the pavement power spectral density is used to reflect different levels of pavement roughness, which is given as follows Gq n\u00f0 \u00de=Gq n0\u00f0 \u00de nf n0 W \u00f05a\u00de where nf is the space frequency (m21), n0 is the reference frequency (its value is 0.1m21), Gq(n0) is a pavement roughness factor (m3), and W is the frequency index. Generally, the value of W is 2. Because frequency O = 2pn, equation (5a) can be rewritten as Gq O\u00f0 \u00de=2pw+1Gq n0\u00f0 \u00de O n0 W \u00f05b\u00de When O! 0,Gq(O)! \u2018, the above equation can be written as Gq O\u00f0 \u00de=2pw+1Gq n0\u00f0 \u00de O+Oc n0 W \u00f05c\u00de According to the random vibration theory, the following relationship can be obtained Gq O\u00f0 \u00de= H O\u00f0 \u00dej j2Sw \u00f05d\u00de Through equations (5c) and (5d) and the power spectral density Sw = 1 of white noise W(s), the spatial frequency response function H(O) can be obtained as follows H O\u00f0 \u00de= n0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2pGq n0\u00f0 \u00de p Oc + jO \u00f05e\u00de By equation (5e), the differential equation of road roughness of front wheels can be deduced dqf s\u00f0 \u00de ds +Ocqf s\u00f0 \u00de= n0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2pGq n0\u00f0 \u00de q W s\u00f0 \u00de \u00f05f\u00de Equation (5f) is the stationary process of road roughness in the spatial domain, as follows dqf s\u00f0 \u00de ds = 1 _sc dqf t\u00f0 \u00de dt \u00f05g\u00de When the vehicle is traveling at a non-uniform speed, equation (5g) is substituted into equation (5f), and the following expression can be obtained _qf(t)+ _scOcqf(t)= _scn0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2pGq(n0) p W s(t)\u00bd \u00f05h\u00de However, as W\u00bds(t) is a non-stationary white noise, it cannot be directly used. The covariance of W\u00bds(t) is E W s(t1)\u00bd W s(t2)\u00bd f g= d s t2\u00f0 \u00de s t1\u00f0 \u00de\u00bd = d t2 t1\u00f0 \u00de _s t1\u00f0 \u00de \u00f05i\u00de Then, the stationary white noise W1(t) is used to define the general non-stationary process W1(t)= ffiffi _s p , and the non-stationary process can be expressed as E W1 t1\u00f0 \u00deffiffiffiffiffiffiffiffiffi _s t1\u00f0 \u00de p W1 t2\u00f0 \u00deffiffiffiffiffiffiffiffiffi _s t2\u00f0 \u00de p \" # = d t2 t1\u00f0 \u00de _s t1\u00f0 \u00de \u00f05j\u00de Through equations (5i) and (5j), we can know that the random processes W\u00bds(t) and W1(t)= ffiffiffiffi _sc p have the same covariance, so equation (5h) can be rewritten as equation (6). In our model, the driving speed of the vehicle is not constant, and the front and rear wheels are differently excited by the road surface. The front wheel road excitation is expressed as _qf(t)+ _scOcqf(t)= n0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2pGq(n0)sc p W1(t) \u00f06\u00de There is a hysteresis between the front and rear wheel road excitations due to the distance between wheels. The rear wheel road excitation is expressed as qr(t)= qf(t t0) \u00f07\u00de From equations (6) and (7), the rear wheel road excitation can be expressed as qr(t) = 2 _sc lc _qr(t) _qf(t)+ 2 _sc lc _qf(t) \u00f08\u00de where sc is the longitudinal displacement of the front wheel, Oc represents the cutoff angular frequency of pavement space (Oc =2pnc), and W1(t) denotes the white noise with zero arithmetic mean. Engine excitation including forces and moments generated by the engine is also a major factor affecting the stability of vehicles. In our model, the vertical force Fz and the moment Mx, generated by a four-cylinder inline four-stroke engine that is commonly used in construction vehicles, are given in equations (9a) and (9b) Fz =4mrl v 2pnR 2 cos v pnR t \u00f09a\u00de Mx =4 Mx 1 2mr2 v 2pnR 2 sin v pnR t mr2 v 2pnR 2 l 2 2 sin 2v pnR t + a2 sin v pnR t+u2 + a4 sin 2v pnR t+u4 + \" # \u00f09b\u00de where M is the average gas moment of the piston, m is the equivalent mass of the round-trip motion, r is the rotation radius of crank, n is the variable ratio, R is the wheel radius, l is the ratio of crank length to the length of connecting rod, v is the vehicle speed, ai is the amplitude of overturning moment, and ui is the phase angle of overturning moment. In our model, the front and rear bodies have different driving and rotation speeds. Considering the speed difference of the front and rear bodies, the longitudinal, lateral, and vertical absolute speeds of the rear body centroid can be expressed as follows _ux2 = _ux1 cosu _uy1 lf2 _c1 sinu \u00f010a\u00de _uy2 = _ux1 sinu+ _uy1 lf2 _c1 cosu cos (Df) _uz1 + lf2 _u1 sin (Df) lr2 _c2 \u00f010b\u00de _uz2 = _ux1 sinu+ _uy1 lf2 _c1 cosu sin (Df) + _uz1 + lf2 _u1) cos (Df + lr2 _c2 \u00f010c\u00de where Df denotes the relative roll angle between the front and rear bodies (Df=f1 f2). Because a suspension system is considered in our model, when road surface excitation forces are transmitted to the suspension system through tires, the relative roll angles of the front and rear bodies will change with time. This change will affect the speed of the rear body in three directions (x, y, and z). In order to facilitate research, Df is simplified to be proportional to f1 (Df= kf1). Correspondingly, the above speed formulas for the vehicle rear body can be changed into _ux2 = _ux1 cosu _uy1 lf2 _c1 sinu \u00f011a\u00de _uy2 = _ux1 sinu+ _uy1 lf2 _c1 cosu cos (kf1) _uz1 + lf2 _u1 sin (kf1) lr2 _c2 \u00f011b\u00de _uz2 = _ux1 sinu+ _uy1 lf2 _c1 cosu sin (kf1) + _uz1 + lf2 _u1 cos (kf1)+ lr2 _c2 \u00f011c\u00de Due to the relative rotation of the front and rear bodies, the relationship between the front and rear body yaw angles is given as follows u=c1 c2 \u00f012\u00de Through equations (11a)\u2013(11c), the longitudinal, lateral, and vertical accelerations of the rear body centroid can be expressed as follows \u20acux2 = \u20acux1 _u _uy1 + _ulf2 _c1 cosu \u20acuy1 lf2\u20acc1 + _u _ux1 sinu \u00f013a\u00de \u20acuy2 = \u20acux1 _u( _uy1 + lf2 _c1 sinu+ \u20acuy1 lf2\u20acc1 + _u _ux1 cosu k _f _uz1 + lf2 _u1 cos (kf1) k _f( _ux1 sinu+ _uy1 lf2 _c1 cosu+ \u20acuz1 + lf2\u20acu1 sin (kf1) lr2\u20acc2 \u00f013b\u00de \u20acuy2 = \u20acux1 _u( _uy1 + lf2 _c1 sinu+ \u20acuy1 lf2\u20acc1 + _u _ux1 cosu k _f _uz1 + lf2 _u1 sin (kf1)+ k _f( _ux1 sinu+ _uy1 lf2 _c1 cosu+ \u20acuz1 + lf2\u20acu1 cos (kf1)+ lr2\u20acc2 \u00f013c\u00de In this paper, the vertical displacement of the suspension system is used as the state variable to analyze the force equations, and we are ultimately concerned with the vertical vibrations ux1 and ux2 of the front and rear bodies, respectively, as shown in Figure 3. Therefore, the vertical displacements of four suspensions need state transition. When the front and rear body roll angles f1 and f2 are small, the centroid displacement in the x direction and pitch angle of the front body can be expressed as uz1 = (z1 + z2)=2 and tanf1 =f1 = (z1 z2)=0:5B1, respectively. Similarly, vertical displacements of the four suspensions are obtained as follows z1 = uz1 +0:25B1 tanf1 \u00f014a\u00de z2 = uz2 0:25B1 tanf1 \u00f014b\u00de z3 = uz3 +0:25B2 tanf2 \u00f014c\u00de z4 = uz4 0:25B2 tanf2 \u00f014d\u00de Substituting equations (13) and (14) into the dynamic equation (4) and eliminating Cx, Cy, and Cx related to the hinge, the equations of motion for a 16- DOF articulated loader can be obtained as follows." + ] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.12-1.png", + "caption": "Figure 1.12. Car trying to detect whether the point m is braking", + "texts": [ + " A higher-order integration scheme, such as that of Runge Kutta, should be used. This is given by: x(t+ dt) . x(t) + dt \u00b7 \" f (x(t),u(t)) 4 + 3 4 f(x(t) + 2dt 3 f(x(t),u(t)),u(t+ 2 3dt)) # . For an easier graphical representation, we will take r = 10 m, *1 = 1 m and g = 10 ms\"2. Discuss the results. EXERCISE 1.8.\u2013 Brake detector We will now look at a problem that involves basis changes and rotation matrices. A car is preceded by another car m (which we will assume to be a point). We attach to this car the coordinate system R1 : (o1, i1, j1) as represented in Figure 1.12. The coordinate system R0 : (o0, i0, j0) is a ground frame assumed to be fixed. This car is equipped with the following sensors: \u2013 several odometers placed on the rear wheels, allowing us to measure the speed v of the center of the rear axle; \u2013 a gyro giving the angular speed of the car !\u0307, as well as the angular acceleration !\u0308; \u2013 an accelerometer placed at o1, allowing us to measure the acceleration vector (\",#) of o1 expressed in the coordinate system R1; \u2013 using two radars placed at the front, our car is capable of (indirectly) measuring the coordinates (a1, b1) of the point m in the coordinate system R1 as well as the first two derivatives - a\u03071, b\u03071 ", + "\u2013 (Floating wheel) 1) Considering the Euler\u2019s rotation equation I!\u0307r + !r ' (I!r) = \" r where the torque \" r = 0, and the fact that we have no acceleration of our wheel, from [1.12] we get ( 999999) 999999* p\u0307 = R (&, !,') \u00b7 vr (i)$ % &\u0307 !\u0307 '\u0307 & ' = $ % 1 tan ! sin& tan ! cos& 0 cos& \" sin& 0 sin% cos $ cos% cos $ & ' \u00b7 !r (ii) v\u0307r = \"!r ' vr (iii) !\u0307r = \"I\"1 \u00b7 (!r ' (I \u00b7 !r)) (iv) [1.16] 2) For the simulation, we take x = (p,&, !,',vr,!r) = (0, 0, 22 34 5 p , 0, 0, 02 34 5 %,$,# , 5, 0, 02 34 5 vr , 5, 1, 02 34 5 !r ) and we get the result depicted in Figure 1.12. The wheel translates with respect to px. We see from the shadow (black) that the rotation axis oscillates. This corresponds to the precession. 3) We have L = RIRT! = RI!r. Therefore L\u0307 = R\u0307I!r +R \u02d9\u00b7I\u00b72345 =0 !r +R \u00b7 I \u00b7 !\u0307r2 34 5 =\"I!1\u00b7(!r$(I\u00b7!r)) = R\u0307I!r \"R (!r ' (I \u00b7 !r)) = R\u0307I!r \"RRTR\u0307I \u00b7 !r = R\u0307I!r \" R\u0307I \u00b7 !r = 0. Moreover E\u0307K = d dt / 1 2L TR!r 0 = 1 2 L\u0307T 2345 =0 \u00b7R!r + 1 2 LT 2345 =!T rIR T \u00b7 - R\u0307!r +R!\u0307r . = 1 2 (! T rIR TR\u0307 \u00b7 !r2 34 5 =0 + !T rIR TR( =!\u0307r4 52 3 \"I\"1 \u00b7 (!r ' (I \u00b7 !r))) See [1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002099_icar46387.2019.8981581-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002099_icar46387.2019.8981581-Figure8-1.png", + "caption": "Fig. 8: Definition of a Multi-Surface Admittance Control to ensures a surface-surface contact between the structuralframe and the panel", + "texts": [ + " During the assembly of the presented parts, the contact forces between them have to be normal and uniformly distributed over all surfaces. Moreover, the process has to respect a maximum normal force at each surface of 100N, and at the end of the process, the gap between the surfaces has to be less than 0.3mm. Assuming all the contact surfaces match perfectly between the panel and the structural-frame, their full contact can be guaranteed if their top and bottom surfaces are in contact. Therefore, the Multi-Surface Admittance Control approach can be implemented at these two surfaces to ensure the desired contact forces. Figure 8 shows the definition of the coordinate systems U1xyz and U2xyz for the bottom and top surfaces, respectively. The robot\u2019s TCP is placed at the origin of the force sensor\u2019s Coordinate system (Sxyz) and follows the same orientation. One admittance controller is defined at each Uixyz with its own controller constants (MdUi , DdUi , KdUi ). The calculation of fd1 and fd2 was made assuming that, at the end of the task, a normal force of 50N is applied at each surface. Each force (fd) was calculated projecting the resultant spatial force measured at Sxyz, using the equations (6) and (7)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.19-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.19-1.png", + "caption": "FIGURE 7.19", + "texts": [ + " The volume of the selected droplet in the experiment is 1 \u03bcL, and the solution can obtain that the droplet radius r0 is 0.62 mm, under which the Bond number is ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi g\u03c1r02=\u03b3lv p 0:23, 1 (7.22) Contact angle hysteresis model. Contact angle changes with droplet volume. where liquid gas interfacial tension \u03b3lv5 71.423 1023 N/m, and water density \u03c15 996 kg/m3. Because the bond numbers are less than 1, we can ignore the gravity of the droplet [14]. The force of liquid drops on a solid surface under the action of airflow is shown in Fig. 7.19. The driving force of a liquid droplet on a solid surface is the component of the driving force of airflow along the direction of the solid surface, and the viscous resistance between a droplet and a solid surface is the delayed friction caused by contact angle hysteresis. They work together to block the movement of the liquid. When the driving force is greater than the resistance of the droplet, the droplet moves on the solid surface. The driving force of the droplet starting is the thrust generated by the airflow along the surface direction of the Fq component" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000019_3305275.3305335-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000019_3305275.3305335-Figure2-1.png", + "caption": "Figure 2. Composition of ADAM", + "texts": [ + " To copy otherwise, or republish, to post on servers or to redistribute to lists, requires prior specific permission and/or a fee. ISBDAI '18, December 29\u201330, 2018, Hong Kong, Hong Kong \u00a9 2018 Association for Computing Machinery. ACM ISBN 978-1-4503-6570-3/18/12\u2026$15.00 DOI:https://doi.org/10.1145/3305275.3305335 2.1 The Composition and the Finite Element Model Parameters The ADAM consists of a series of identical span units, each of which is composed of transverse bar, longitudinal bar, spherical joint, stay cable component and guide wheel, as shown in Figure 2. Transverse bar and longitudinal bar are connected through spherical joint, and guide wheel ensures the ADAM to expand smoothly, while stay cable component provides tension force for the mast. According to actual application, the geometrical dimensions of transverse bar and longitudinal bar are shown in Table.1. ANSYS a high-efficiency finite element analysis software integrating structure, fluid, heat, magnetic, etc., and applicable to the ADAM described in this paper, which can improve the efficiency of simulation greatly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002380_j.mechmachtheory.2020.103858-Figure16-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002380_j.mechmachtheory.2020.103858-Figure16-1.png", + "caption": "Fig. 16. Case: 2 - Initial posture of the ball w.r.t. bowl.", + "texts": [ + " In this figure, it can be observed that after an initial spike the normal reaction force approaches the weight of the ball. Fig. 15 (c) shows the magnitude of the resultant frictional force on the ball, which is a spike of very small magnitude of the order of 10 \u22127 , and has already been explained in the previous subsection. The animation for this subcase has been shown in the video attached as supplementary file \u2018 Case1 Subcase2 Ball Cartilage Bowl.avi \u2019. In this case, the ball is dropped vertically on the bowl with an offset of 100 mm along the positive X -direction of the inertial frame { 0 }, as shown in Fig. 16 . The initial height of the COM of the ball is 150 mm along the positive Y -direction of the inertial frame { 0 }. The values of constants K in (21) and R in (23) are 10 and 100 respectively for this case. Fig. 17 shows the loci of the COM of the ball C B and position of a point B i on its surface for Case 2. The respective loci of these two points are shown in purple and orange colors. A line segment C B B i joining these two points is shown in gray. It can be observed that these loci begin as straight vertical segments until the ball contacts the cartilage layer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003698_rusautocon49822.2020.9208181-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003698_rusautocon49822.2020.9208181-Figure1-1.png", + "caption": "Fig. 1. The appearance of the mining shearer-loader", + "texts": [ + " Keywords\u2014shearer-loader, forward motion drive, cutting drive, mathematical model, transients, coal strength, control system I. INTRODUCTION One of the areas of the technical process in the coal industry is the widespread introduction of systems and means of automation of mining equipment and, first of all, shearers. Existing mining shearer-loaders extracting minerals (coal, silvinites) have a cutting drive and a forward motion drive which interact among themselves through an executive body for destruction of a layer of a mineral - Fig. 1 [1, 2]. A coal seam is a heterogeneous mass with the presence of solid inclusions of rock, which, when destroyed, leads to strong shock loads on the executive body of the shearer. Thus, the cutting drive is one of the weakest part of the shearer [3]. Automation of operations of the shearer should increase their productivity due to a more complete use of the energy capabilities of the electric drive and reduce a likelihood of \"induction motor stalling\", increase durability due to reduced overloads of the electric motor and the mechanical part, and hence their accident rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002950_042044-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002950_042044-Figure6-1.png", + "caption": "Fig 6. Temperature field distribution of permanent magnet coupler.", + "texts": [ + "5, and the heat dissipation coefficient is calculated according to the linear speed of each part, unit is W/(m2\u00b7K). Simplifying the three-dimensional model of the permanent magnet coupling to established the finite element analysis mode, as shown in the Fig. 5. The conductor cylinder and the outer steel plate on the same side are set as the thermal load, assuming that the temperature distribution of the conductor cylinder and the outer steel plate is uniform, add internal h load to each generation and the surface heat transfer coefficient are meshed and analysed. Fig. 6 is the simulation analysis result of temperature field of cylinder permanent magnet prototype, and under 20rpm normal temperature. It can be seen from Fig.6 that the temperature of conductor cylinder, inner aluminium yoke and permanent magnet is higher, and the temperature of outer aluminium yoke, permanent magnet and steel frame is lower. This paper makes an electromagnetic field simulated analysis on the cylindrical permanent magnet coupling, the result comes out the largest area of the conductor tube eddy current density, and the largest area of the density of the eddy current loss. The simulation analysis on temperature field of the IWAACE 2020 Journal of Physics: Conference Series 1550 (2020) 042044 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001291_aim.2019.8868674-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001291_aim.2019.8868674-Figure6-1.png", + "caption": "Fig. 6. Test setup of dual-mass drive system with backlash and compliance.", + "texts": [ + " In order to suppress the limit cycle caused by the backlash nonlinearity, the eigenvalue curve over the entire amplitude must be below the zero axis. To achieve this, the gain value Kp\u03c9 of the PI speed controller is reduced to Kp\u03c9 = Kcrit = 0.063 until there is no intersection with the zero axis (see fig. 5). This statement will be confirmed experimentally in chapter IV. The proposed limit cycle analysis for the dual-mass system with backlash nonlinearity is validated, using the experimental setup depicted in Fig. 6. It comprises two synchronous machines. The left synchronous machine serves as a drive motor. The right motor can be used to simulate the influence of disturbance as load torques, which is not dealt with in this paper. The current control of the synchronous machines is implemented on their frequency converters. The current controller is designed according to the magnitude optimum criterion as described in chapter II. For position and speed measurements on the motor side, a high-resolution incremental encoder with sinusoidal track signals is used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001441_012149-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001441_012149-Figure8-1.png", + "caption": "Figure 8. 2-pole eLNG synchronous motor. Figure 9. 2-pole eLNG induction motor.", + "texts": [ + "64 MTPA each. Each train is composed by three motor driven refrigerant compression lines with 75 MW shaft power. From the process point of view, each string is designed with three large single compressors (Propane, Low Pressure LP Mixed Refrigerant, Medium/High Pressure M/HP Mixed Refrigerant) on three independent shaft lines driven by a large 2-pole synchronous motor fed by the 60Hz public electric network through a Load Commutated Inverter (LCI) system requested by the customer (see fig.7 and fig.8). The speed is 3000 rpm to avoid any risk of inter-harmonic coupling with the grid and torque pulsations with the shaft-line. The selection of motor as prime mover is driven mainly by necessity to contain dimension and weight to better adapt to plant characteristics. With respect to a traditional solution driven by gas turbine, an electric motor of same power size allows a length reduction to approximately half time and consistent weight reduction and the elimination of air filter chamber of the gas turbine [11]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003628_j.procir.2020.09.063-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003628_j.procir.2020.09.063-Figure2-1.png", + "caption": "Fig. 2. Two additive nozzle designs integrated into the processing head", + "texts": [ + " In combination, it results in direct metal deposition onto a substrate (see figure 1). With the help of a dichroitic mirror, a camera can observe the melt pool and the deposition during the process. In this paper, different concepts of the powder supply system are evaluated and compared to the conventional lathemanufactured approach fitted to the machine. In order to maintain cross compatibility with the processing head and ensure comparability of the new designs, the nozzles are adapted to an existing process head (see figure 2). In consequence, the maximum radial size of the nozzle is restricted to 24 mm. Within these constraints, the inner-structure of the nozzle can be modified freely by using the advantages of laser powder bed fusion additive manufacturing and the resulting design freedom. The AM nozzles themselves are attached with a built-in thread to the mount on the process head. To ensure exact positioning and overcome thread tolerances, both parts are centered by two conical mating surfaces that are integrated into the assembly", + " They are an evolution of conventional nozzles with annular openings. The outlet gap sizes are varied from 0.2 mm to 1.0 mm (see table 1). Furthermore, the use of powder bed additive manufacturing allows for asymmetric designs with undercuts. Therefore, a honeycomb design with a complex inner structure and 0.4 mm wide channels is created and evaluated. The process head has integrated cooling and powder supply connections. With help of the conical mating surfaces, the AM nozzles are centered in the thread coaxially to the laser beam. In figure 2, schematic cross sections of two additive designs are shown. A nozzle using an annular outlet is visible on the bottom left. On the bottom right, individual channels of the honeycomb design can be seen. According to table 1, one conventional and five additively manufactured nozzles are evaluated. The AM nozzles are created by laser powder bed fusion (LPBF) using CuSn10 material. This tin bronze is used for better heat conduction compared to commonly used 316L stainless steel. The manufacturing was carried out on an Orlas Creator LPBF machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001298_aim.2019.8868514-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001298_aim.2019.8868514-Figure5-1.png", + "caption": "Fig. 5. A course for testing the performance of running with dragging the cable", + "texts": [ + " 2) the course is for moving around the objects and it includes turning left section and turning right section. \u2022 Metrics: 1) the test shall be conducted by 10 repetitive round- trips and in 30 minutes 2) keep the records of measured total time and calculated average of moving speed based on a predetermined reference distance of the course. 3) confirm the whole robot body completely reaches start point zone and halfway point zone of the course. \u2022 An example of practical course : Upper figure of Figure 5 shows a typical example of the test course that satisfies above-mentioned conditions. Bottom figure of Figure 5 shows a single round-trip route for testing. Figure 6 shows the pillar type object (planar size: 1.2 m \u00d7 1.2m) which consists of three mutually perpendicular cuboids. The design concept of this object is to create a structure that provide the hazards of easy entwinement of the cable and the load effect for dragging the cable. In this case, two pillar type objects are placed back-to-back and there is 1.0m interval. The robot moves according to the numbers and after reaching at the halfway point, it moves according to the reverse order" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000530_978-3-030-12082-5_55-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000530_978-3-030-12082-5_55-Figure10-1.png", + "caption": "Fig. 10 Distribution of temperatures in gyro unit-platform assembly model, \u00b0C", + "texts": [ + " The thermal contact conductance is significantly affected by clamp force of the bolts. Experiment No. 6 showed that using the conditions of experiment No. 4 and decreasing the clamp force from 2000 to 100 N, the averaged thermal contact conductance decreased by more than 6 times. A similar effect in real structures can arise in the case of more complicated connections, for example, with clasps [35]. The distribution of contact pressures, surface temperatures, and heat fluxes for the contact platform surface in experiment No. 4 is shown in Figs. 7, 8, and 9. Figure 10 shows the distribution of temperatures throughout the entire model of the gyro unit-platform assembly under the same conditions. The model problem of the contact of the gyro unit-platform demonstrates the applicability of the developed simulation method for the contact of rough surfaces at the micro-level [11] to macroscale objects. It is determined that essential factors affecting the real contact area and, consequently, the temperature distribution in the macroscale (i.e., curving the surfaces) are both thermal expansion and uneven pressure due to the tightening of the bolts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure50.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure50.2-1.png", + "caption": "Fig. 50.2 a Weld direction and rolling direction, b torch and fixture equipped with SOD adjustment, c welded plate", + "texts": [ + " The setup has been shown in Fig. 50.1. To enhance percentage elongation, the weld direction has been determined with special light scattering technique which always shines in the direction perpendicular to the rolling direction. Due to rolling, grains are elongated in the direction of rolling resulting in less percentage elongation and hardness in the rolling direction. Therefore, care has been taken during the experiment and welding direction has been proposed normal to the light scattering direction [2]. Figure 50.2a illustrates the weld direction and rolling direction with the help of light scattering direction. 50 Multi-output Response Optimization for Overall Enhancement \u2026 589 The yellow arrow shows the rolling direction, and the blue arrow shows the welding direction. The welding setup, fixture, and standoff distance adjustment features for MPAW setup were fabricated at IIT Guwahati central workshop for accuracy of weld quality, which has been shown in Fig. 50.2b, a welded sample was shown in Fig. 50.2c. The weld fixture was also designed as per dimension of weld sheet and less arc deflection. The interface gap between two weld plates, which is not permissible was ensured properly before welding. The chemical composition of the weld material used in these experiments has shown in Table 50.1. The significant six parameters in three levels which have an impact on weld quality, taken into consideration and shown in Table 50.2. During post-weld process, tensile samples were cut through CO2 laser-cutting machine, make: model no: LVD Orion3015, power rated 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003718_icra40945.2020.9197100-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003718_icra40945.2020.9197100-Figure3-1.png", + "caption": "Fig. 3: The links are joined by shoulder bolts that each hold a retract and grip pulley. Cables wrapped in opposite directions provide antagonistic actuation and are both fixed at the tip.", + "texts": [ + " Using two actuators instead of one to control the linkage offers the ability to simultaneously scale the tension in each pulley series. If both the grip and retract tension are scaled by the same factor, the torque pattern is preserved and system stability is maintained. Similarly, the forces into a surface will be scaled by the same ratio, increasing friction. We present a gripper with two opposing linkages similar to that presented by Hirose et al. [11]. The transmission consisted of two independently driven pulley mechanisms that opposed one another to generate a net torque at each joint (Fig. 3). We improved on the prior mechanisms by designing the retract and grip pulleys in tandem to achieve the desired joint torques and adding hard-stops to set a specific retract configuration that prevents hardware damage (retracted position seen in Fig. 3). We also used bearings at each joint to reduce friction. Plain bushings were not used because they develop substantial friction that increased with the joint load. Each link also had mounting points for interchangeable interface tiles that can be outfitted with different materials. The interface tiles had a large impact on grip strength and four different types were evaluated on a variety of surfaces (Section V). The links and pulleys were 3D printed from carbon fiber filled PLA to make them lightweight, stiff, and easy to manufacture" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000607_978-3-030-14907-9_87-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000607_978-3-030-14907-9_87-Figure4-1.png", + "caption": "Fig. 4. Frames vision - manipulation system", + "texts": [ + " Focal length parameter was taken from a reference picture that later was used to compare with subsequent pictures and in this way estimating different distances from camera to object. F = P \u00d7 D W (6) D\u2032 = F \u00d7 W P (7) Subsequent pictures taken after reference picture utilized triangle similarity to determine a new distance (D\u2032) from object to camera. Every time the distance between camera and object was changed, D\u2032 took a new value according to constants (F) and (W) and variable (P) as seen in Eq. 7. Once D was known, named x coordinate, (see Fig. 4) it was necessary to determine the position of object in a coordinate plane, located in the camera of the robot. To do this, it was needed to know the number of pixels of the object and making a comparison among physical world and digital world. In consequence, it was possible to learn the dimension in centimeters of the image, which was useful to know the position of the object over a y,z plane (see Fig. 4). Vision algorithm pseudocode is shown below (see Fig. 3) Cartesian coordinates that define localization of the object were already established. The next step was to relate camera frame and robot\u2019s arm frame in order to generate communication and understanding between vision algorithm and manipulation algorithm. The next matrix (see Eq. 8) show the relation among 3 different frames located at camera, object and arm\u2019s base (see Fig. 3). The angle \u03b2 defined pitch in the head of robot, #\u00bb P I0 = (xI0 , yI0 , zI0) defined position of arm\u2019s end effector, which was a 3D vector, #\u00bb PC = (xC , yC , zC) defined camera position, which was a 3D vector and #\u00bb P I = (xI , yI , zI) defined image position, which was a 3D vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000845_icuas.2019.8797739-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000845_icuas.2019.8797739-Figure1-1.png", + "caption": "Fig. 1. Frame and position definition for multirotor control laws.", + "texts": [ + " Velocity tracking is accomplished by reformulating the input of the path-following control law in terms of the desired velocity and heading. The remainder of this paper is organized as follows: in Section II, we define a model of the multirotor dynamics. The formulations of the rate- and attitude-tracking controllers are presented in Section III, and Section IV describes the path-following and velocity-tracking control laws. Section V presents a path-following simulation example, and Section VI concludes the paper. A simplified mathematical model of the multirotor dynamics is used for control law development. Figure 1 depicts the inertial and body-axis reference frame definitions. Two 978-1-7281-0332-7/19/$31.00 \u00a92019 IEEE 353 equivalent representations will be helpful for design of the rotational and path-following control laws. For design of the rotational control laws, multirotor dynamics are represented as x\u0307(t) = v(t) (1a) mv\u0307(t) = mge3\u2212 f (t)b3(t) (1b) q\u0307(t) = 1 2 Rq(t)\u03c9 (t) (1c) J\u03c9\u0307 (t)+\u03c9 (t)\u00d7J\u03c9 (t) =M(t), (1d) where x(t) and v(t) are the position and velocity, respectively, of the vehicle; m is the vehicle mass; J is the inertia matrix; g is the magnitude of gravitational acceleration; e3 is the inertial z-axis unit vector; f (t) is the net thrust applied by the vehicle; and M is the total moment applied to the vehicle", + " The decoupling of spatial tracking from temporal constraints offered by path-following control is convenient, and vehicle behavior is typically smoother2 with path-following control than with trajectory tracking [20], [13]. The path-following control law assumes the existence of an inner-loop controller that can provide adequate tracking of angular rate commands. Exact rate tracking is not required, but bounds on path-following error are dependent on the inner-loop rate tracking performance. Let I = {e1,e2,e3} denote an inertial reference frame, and B = {b1,b2,b3} denote the multirotor body frame, as depicted in Fig. 1. Letting xd(\u03b6 (t)) be the twicecontinuously-differentiable desired path parameterized by a continuously differentiable timing function \u03b6 , and letting x(t) denote the inertial position of the multirotor, we can define the position and velocity errors as ex(t) = xd(\u03b6 (t))\u2212 x(t), (16) ev(t) = x\u0307d(\u03b6 (t))\u2212 x\u0307(t) = dxd(\u03b6 (t)) d\u03b6 \u03b6\u0307 (t)\u2212 x\u0307(t). (17) We define \u03b8 (t) = \u03b6\u0307 (t) (18) to be the mission progression rate, a free variable which can be modified online for time coordination of multiple vehicles [21] or for collision avoidance [22]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003730_tmag.2020.3027291-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003730_tmag.2020.3027291-Figure2-1.png", + "caption": "Fig. 2. Magnetic field simulation of PMTM. (a) Relative position of each part of PMTM. (b) The magnetic lines of the planet gear and central worm stator.", + "texts": [ + " When the planet gears revolve, it can be regarded that the relative position of the permanent magnet teeth and the central worm stator does not change in the range of the wrap angle, so there is no end effect. Therefore, the following analysis of the end-effect force is mainly aimed at the motion of the rotation of the planet gears. In the torque fluctuation analysis of PMTM, the end torque and cogging torque always exist when the planet gears rotate. In order to avoid the influence of cogging torque and electromagnetic torque harmonics, a finite element analysis model of the end-effect force for PMTM is built. The magnetic field simulation of PMTM is shown in Fig. 2. Fig. 2(a) shows the relative position relationship of the outer stator, planet gears and central worm stator. The distribution of the magnetic lines of the planet gear and central worm stator is shown in Fig. 2(b). Within the wrap angle, the magnetic field lines of the permanent magnet teeth form a closed magnetic circuit with the adjacent permanent magnet teeth through the air gap and the central worm stator core. The magnetic field lines of the permanent magnet teeth pass through the air to form a closed magnetic circuit with adjacent permanent magnet teeth when outside the wrap angle, but their magnetic field lines are relatively sparse. The magnetic induction intensity of the end magnetic field of the central worm stator changes abruptly when the permanent magnet teeth enter and exit the wrap angle, because the magnetic permeability of the central worm stator is much larger than that of air" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002856_012016-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002856_012016-Figure2-1.png", + "caption": "Figure 2. Schematic diagram of manipulator model.", + "texts": [ + " From the above modeling process, the general formula of the transformation matrix of the adjacent rod coordinate system is: (Note: () cos(), () sin()C S ) l l 0 l l l l l l l l l l l l C S S S C C Q (5) Wherein i i i iC S T1= =l l lz l l l lz l l l Q Q Q Q Q (6) The rotational transformation matrix l l Q of the D-H system is a unit orthogonal matrix, and its inverse matrix l l Q is equal to its own transpose. The characteristic of l l Q provides great convenience for solving the inverse kinematics of manipulator. 3. Kinematics analysis of the manipulator 3.1. Forward kinematics of the manipulator The schematic diagram of the industrial robot arm is shown in figure 2. From bottom to top, it consists of joints such as base, waist, shoulder, elbow, and wrist. The kinematics model of the manipulator is obtained by obtaining the position and attitude of the end of the manipulator relative to the base of the manipulator. D-H Frame was established according to the kinematics model of the manipulator, D-H parameters are shown in table 1. According to the D-H Frame of manipulator, the positive equation of motion is: 1 1 1 2 1 2 3 7 2 2 3 2 3 4 1 3 4 1 4 5 1 5 6 3 4 5 4 5 6 5 6 7 \\ Q Q Q Q Q Q Q Q Q r r r r r r r (7) AAME 2020 IOP Conf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002970_012003-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002970_012003-Figure2-1.png", + "caption": "Figure 2. Working zone of electric-discharge milling system: (1) gantry; (2) support; (3) head; (4) pulse generators; (5) bath.", + "texts": [ + " The required value of the interelectrode gap is supported by an electromechanical servo system, by moving the servo head (4) in the vertical direction by means of an actuator motor (5), a worm gearbox (6) and a lead screw (7). The engine (5) is controlled by a servo unit. The performance of alloying is determined by the speed of movement of the alloying electrode relative to the hardened surface. The equipment for Electrical Discharge Machining (EDM) \u201cELFA731\u201d provides the movement of the alloying electrode along the coordinates X, Y, Z (figure 2). \u201cELFA-731\u201d machine is equipped with a CNC system \u201cFanuc- 3M\u201d. This CNC device is focused on manual input of control information in a dialogue mode. Due to this, it is possible to prepare a control program directly at the machine, based on the drawing data. Surface reinforcement consists in hardening local areas evenly spaced on the surface. Such processing can provide the necessary level of operational properties with a significant increase in productivity. In addition, surface reinforcement allows you to combine the properties of the base material and the material of the hardened sections" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002942_scc47175.2019.9116173-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002942_scc47175.2019.9116173-Figure3-1.png", + "caption": "Fig. 3: Two-mass-model System.", + "texts": [ + " In the above part, the architecture and performance of the algorithm are well developed, while the emphasis will be placed in the next part on the choice of the numerical example of the dynamic system. III. NUMERICAL EXAMPLE In this section, we will give a numerical example to show the effectiveness of the novel structure based on Lyapunov stability. Throughout this section, the training signals are generated offline with a sampling period Ts = 10ms and the initial coefficients of the network in offline training are randomly initialized. We consider the two-mass electromechanical system model in Fig.3. The two-mass system is a simplified model of an electromechanical system. It is composed of three parts: motor inertia, load inertia, and a low-stiffness shaft. Cm: Motor Torque; \u03c9m: Motor Speed; \u03c9c: Load Speed; \u03b1: Damping Coefficient; K: Shaft Stifness; Cc: Load Torque; Using the Lagrange formalism, the dynamics of the two-mass system can be described by the following dynamic equations [10], [11], [12]: Jm d\u03c9m dt = Cm \u2212 Ct (9) Jc d\u03c9c dt = Ct \u2212 Cc (10) Ct = K \u222b (\u03c9m \u2212 wc) dt+ \u03b1 (\u03c9m \u2212 \u03c9c) (11) With Jm: Motor inertia; Jc: Load inertia; Ct: Shaft torque" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001748_s1052618819060074-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001748_s1052618819060074-Figure3-1.png", + "caption": "Fig. 3. Cinematic chains of the mechanism: (a) single PUS chain and (b) double PUS chain with parallelogram.", + "texts": [ + " Under parametric synthesis of any parallel mechanism, it is necessary to investigate the possibility of special positions arising [10\u201313], as when the mechanisms ingress into these positions, the mobility of 503 output link is changed (the loss of a degree of freedom or uncontrollability arise) [14]. This work is dedicated to solving this problem for a mechanism with four degrees of freedom. For analysis of the specific positions of any parallel mechanism, it is necessary firstly to solve the reverse problem on the positions that can form analytically by analyzing the structure of mechanism and solving the constraint equation relative to the input coordinates [15]. The mechanism under consideration includes kinematic chains of two types: a single PUS-chain (Fig. 3a) and a double PUS-chain (Fig. 3b) with a parallelogram. The position of the output link is specified by coordinates x, y, z of point E in the system Oxyz, and the orientation is specified by the angle of rotation \u03d5y relative to axis y. The input coordinate of the chain is h = zA, limited by the limits of hmin and hmax. If in some initial position the output link is located horizontally and the points E and D are located strictly above the point O, then under known geometric sizes of the chain, the coordinates , , of points C in the initial position in the system Ex'y'z' can be considered specified" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002394_sii46433.2020.9026255-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002394_sii46433.2020.9026255-Figure3-1.png", + "caption": "Fig. 3. Coordinates of the acceleration sensor and gyroscope.", + "texts": [ + " Moreover, RT.1 and 2 prevent falling down and the tendency of turning downhill on a cross slope to adapt to various road surfaces. The power of assist and braking and the limitation of speed can be adjusted. RT.2 has a handbrake, and the users can intentionally stop RT.2 using the handbrake. RT.1 and 2 have several sensors and a data communication module; therefore, the state of the robot can be quantified, and RT.1 and 2 can upload these sensor data to the data server. The robot system is shown in Fig. 3(c). RT.1 and 2 have a two-axis (x and y axes) acceleration sensor, threeaxis gyroscope sensor, GPS sensor, and touch sensor on the handle. The speed of the left and right wheels and current of the left and right motors can be measured. Only RT.1 has force sensors for the handle. The coordinates of the sensors are shown in Fig. 3. RT.1 and 2 have a cellular device and can upload the sensor data via a 3G/LTE base station to the Internet, and these sensor data are stored in the cloud server. In this study, Amazon Web Service (AWS) is used as the cloud storage server. RT.1 and 2 measure the acceleration, gyro, touch, speed of wheel, and current of motor every 1 s and upload them every 1 min. Moreover, RT.1 and 2 measure the GPS information every 1 min and upload the GPS information and the state of the robot (powered, battery remaining, ID of robot, product type) every 10 min" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure51.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure51.4-1.png", + "caption": "Fig. 51.4 Extracting the xyz coordinates from CAD geometry at three sections (top, middle, and bottom)", + "texts": [ + " The nanotom CT works at a tube voltage of 110 kV and current of 80 lA, with a minimum voxel size of <500 nm. The array of 3 3 microholes is scanned, and the scanning time is approximately 45 min. From the generated volume data file, volume graphics (VG viewer) open-source software creates 3D and 2D images for visualization [7]. The 2D and 3D views of the component are shown in Fig. 51.3. In the present work, GOM Inspect open-source software [8] is used in the next step to generate the geometric model as shown in Fig. 51.4. Computed tomography volume data (.vgl format) file is directly imported in GOM Inspect software, without converting into STL format. This is an easy 51 Characterization of Geometrical Features of Ultra-Short Pulse \u2026 605 approach for the dimension analysis of CT-scanned complex shape and size. After importing CT volume data in GOM inspect software, it converts into mesh elements. Cylinder surfaces are constructed by selecting the holes. The coordinate transformation is done by using axis alignment with respect to z-axis in vertical direction. From the geometric model, dimensional and form errors of microholes are assessed as outlined in Sect. 51.4. 606 K. Kiran Kumar et al. 51 Characterization of Geometrical Features of Ultra-Short Pulse \u2026 607 To compare with the results obtained from GOM Inspect software, least squares technique developed by the authors is used for evaluation of form errors. The coordinates of microholes are extracted using GOM Inspect at selected sections as shown in Fig. 51.4. The xyz coordinates are extracted at top, middle, and bottom sections, respectively. The xyz coordinates at top surface of one microhole are given in Table 51.1. The extracted coordinate data of microholes contains both size and form information, and a suitable method has to be implemented to separate the form data and size information. Samuel and Shunmugam [9] have proposed a transformation technique for evaluation of form error from the coordinate data. Accordingly, a reference circle is chosen passing through three selected points and the deviations are computed using Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-Figure12-1.png", + "caption": "Fig. 12. Integrated experiment device", + "texts": [ + " In this experiment, the tip position was controlled with a resolution of 50 mm by bending. In order to achieve this control only by the rotation of the base, a resolution of 0.84 deg is required. In addition, if the entire arm is moved by the rotation of the base, it is difficult to stabilize the tip position because the arm vibrates. Therefore, the bending motion is more suitable for controlling the tip position. In this section, we integrated the linear mechanism and bending mechanism and describe the experiment on the combined motion of linear motion and bending motion. Fig. 12 shows an overview of the experiment device that integrates the linear mechanism and the bending mechanism. The bending mechanism and the telescopic structure are connected to the linear mechanism. The telescopic structure has 7 nodes and a total length of 2080 mm, and one rope guide with a pulley is arranged. We performed an experiment to move linearly while bending as the combined motion of linear motion and bending motion. This is the operation necessary to achieve linear motion while avoiding obstacles on the axis of extension" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003172_042002-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003172_042002-Figure2-1.png", + "caption": "Figure 2. Vehicle workspace in X,Y,Z plane", + "texts": [ + " Therefore, some researchers are working to build a mathematical model of vehicle movement and its environment. Based on the above design, the vehicle model can be derived by looking at the vehicle as a point in the 3 dimensional plane of cartesius coordinates with global coordinates {X, Y, Z}. The vehicle has 5 (five) position parameters represented as pc = {xc, yc, zc, c, c}, where (xc, yc, zc) is the spatial position of the vehicle in the global coordinate system and ( c, c) are respectively each vehicle\u2019s direction angle to the X axis and Z axis. This principle is shown in Figure 2 below. Journal of Physics: Conference Series 1569 (2020) 042002 IOP Publishing doi:10.1088/1742-6596/1569/4/042002 The important thing to note in this vehicle movement model as previously explained is the 4 degrees of freedom of motion as shown in Figure 2. Thus the vehicle can be modeled by observing dynamic parameters nonlinearly, coupled, time-varying, and uncertainty. [15] model the parameters as follows. MRXb + C(Xb)Xb + D(Xb)Xb + g(XE) = F + Fd (1) F = [X Y Z K M N]T (2) Xb = [u v w p q r]T (3) where MR is matrix of vehicle body mass movement and water mass; C is matrix of centripetal force and Coriolis; D is matrix of hydrodynamic dumping; G is matrix force and inertia moment. The body frame of the vehicle moves and always has a normal position at the center of gravity. From equations (1), (2) and (3), Xb is the velocity vector of the linear state of the vehicle body (surge, sway and up) and angular (roll, yaw and pitch) as shown in Figure 2. F is a force vector and torque produced by the degree of freedom of motion and Fd are vectors that represent environmental disturbances. Other parameters are MR, C (Xb), D (Xb) and g (XE) defined as follows, (4) (5) (6) In other conditions, vector XE = [x, y, z. \u03c6, \u03b8, \u03c8] are variable conditions; state of earth space: movement of the position of x, y and z and yaw angle, up and roll. The relationship between the speed of the vehicle's body and the state of the earth's space can be written as follows, XE = J(XE)Xb (7) Based on the above design, there are 2 (two) pairs of motors that will produce 2 (two) pairs of speeds, namely: the translation speed v1, and v2, and the rotational speed 1 dan 2, where v1 and 1 are motor speed X1 and X2 and v2 and 2 are motor speed Z1 and Z2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure19-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure19-1.png", + "caption": "Figure 19. Displacement distribution from a plain bearing to a textured surface (N = 6000rpm, W = 10000N)", + "texts": [], + "surrounding_texts": [ + "graph\ufeffclearly\ufeffshows\ufeffthat\ufeffthe\ufeffincrease\ufeffin\ufeffthe\ufeffspeed\ufeffof\ufeffrotation\ufeffof\ufeffthe\ufeffshaft\ufeffleads\ufeffincreased\ufeffhydrodynamic\ufeff pressure,\ufeffthe\ufeffsignificant\ufeffvalues\ufeffare\ufeffnoted\ufefffor\ufeffa\ufeffspeed\ufeffof\ufeff9000\ufeffrpm,\ufeff(Figure\ufeff17\ufeffand\ufeffFigure\ufeff18).\nDisplacement Distribution for Textured Plain Bearing The\ufeffmaximum\ufeffdisplacement\ufeffdue\ufeffto\ufeffthe\ufeffhydrodynamic\ufeffpressure\ufefffield,\ufeffis\ufeffnoted\ufefffor\ufeffan\ufeffangular\ufeffposition\ufeff of\ufeff150\ufeff\u00b0\ufeffto\ufeff200\ufeff\u00b0,\ufeffas\ufeffwell\ufeffas\ufeffit\ufeffis\ufeffinfluenced\ufeffby\ufeffthe\ufeffincrease\ufeffof\ufeffthe\ufeffrotational\ufeffvelocity\ufeff(Figure\ufeff19).\ufeff increasing\ufeffthe\ufeffspeed\ufeffof\ufeffrotation\ufeffof\ufeffthe\ufeffshaft\ufeffleads\ufeffto\ufeffthe\ufeffincrease\ufeffin\ufeffdisplacement.\nThe\ufeffmaximum\ufeffvalue\ufeffof\ufeffdisplacement\ufeffis\ufeffnoted\ufefffor\ufeffa\ufeffspeed\ufeffof\ufeff12000rpm\ufeffin\ufeffthe\ufeffcase\ufeffof\ufeffa\ufefftextured\ufeff plain\ufeffbearing\ufeffis\ufeffof\ufeffthe\ufefforder\ufeffof\ufeff0.39\u03bcm.\ufeffOn\ufeffthe\ufeffother\ufeffhand,\ufefffor\ufeffa\ufeffplain\ufeffbearing\ufeffwithout\ufefftexture,\ufeffthe\ufeff maximum\ufeffvalue\ufeffof\ufeffthe\ufeffdisplacement\ufeffis\ufeffof\ufeff0,32\u03bcm,\ufeffFigure\ufeff20.\ufeffThe\ufeffdisplacement\ufeffdistribution\ufeffallowing\ufeff to\ufeffangular\ufeffposition\ufefffor\ufeffdifferent\ufeffrotational\ufeffvelocity\ufeff(6000,\ufeff9000\ufeffand\ufeff12000rpm),\ufefffor\ufeffplain\ufeffbearing\ufeffat\ufeff textured\ufeffsurface\ufeffand\ufeffuntextured\ufeffplain\ufeffbearing\ufeffis\ufeffillustrated\ufeffin\ufeffthe\ufeffFigure\ufeff21.\ufeffthe\ufeffmaximum\ufeffdeformation\ufeff is\ufeffnoted\ufeffat\ufeffthe\ufefflevel\ufeffof\ufeffthe\ufefflower\ufeffgeneratrix,\ufeffas\ufeffwell\ufeffas\ufeffthe\ufeffdisplacement\ufeffvariation\ufeffof\ufeffbearing\ufeffwith\ufeffthe\ufeff rotational\ufeffvelocity\ufeffvariation\ufeffis\ufeffsignificant\ufefffor\ufeffa\ufeffplain\ufeffbearing\ufeffwith\ufeffa\ufefftextured\ufeffsurface\ufeffcompared\ufeffto\ufeffthat\ufeff obtained\ufefffor\ufeffa\ufeffnon-textured\ufeffplain\ufeffbearing.\nShear Stress The\ufeffshear\ufeffstress\ufeffevolution\ufeffaccording\ufeffto\ufeffangular\ufeffposition\ufefffor\ufeffdifferent\ufeffrotational\ufeffvelocity\ufefffor\ufeffplain\ufeff bearing\ufeffat\ufefftextured\ufeffsurface\ufeffand\ufeffuntextured\ufeffplain\ufeffbearing,\ufefffor\ufeffthe\ufeffradial\ufeffload\ufeffis\ufeff6000N,\ufeffis\ufeffillustrated\ufeff in\ufeffthe\ufeffFigure\ufeff18.\ufeffThe\ufeffmaximum\ufeffstresses\ufeffshear\ufeffvalues\ufeffare\ufeffnoted\ufeffin\ufeffthe\ufeffboth\ufeffangular\ufeffpositions\ufeff60\u00b0\ufeffand\ufeff 230\u00b0\ufeff(Figure\ufeff22).\ufeffThe\ufeffincrease\ufeffin\ufeffthe\ufeffrotational\ufeffvelocity\ufeffof\ufeffthe\ufeffshaft\ufeffleads\ufeffto\ufeffan\ufeffincrease\ufeffin\ufeffthe\ufeffshear\ufeff stress,\ufeffFigure\ufeff23.\ufeffThis\ufeffincrease\ufeffreaches\ufeff23\ufeffpercent\ufefffor\ufeffa\ufefftextured\ufeffplain\ufeffbearing,\ufeffon\ufeffthe\ufeffother\ufeffhand\ufefffor\ufeff a\ufeffnon-textured\ufeffplain\ufeffbearing\ufeffreaches\ufeff23\ufeffpercent." + ] + }, + { + "image_filename": "designv11_80_0000542_978-3-030-20131-9_261-Figure65-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000542_978-3-030-20131-9_261-Figure65-1.png", + "caption": "Fig. 65. Schematic diagram of beaded cool pad weaving device.", + "texts": [], + "surrounding_texts": [ + "In chapter 2, three methods of automatic knitting of bead mat are put forward. Among them, Warp and Weft Automatic Weaving Method is the most simple, but the cost is higher and its stability is poor. The lock stitch sewing weaving method is rea lly good, but there exists three difficulties applied in the machine: The cross-sectional size of the hook must be smaller than the size of the bead hole as the hook needle needs to completely pass through the bead, which makes it difficult to hook the string smoothly. It is difficult to guarantee the parallelism between the cross-section of the loop ring and the end of the bead of the first row in steps 4-6. When weaving a larger size beaded cooling pad, threading the string into the longer aligned transverse bead holes becomes hard. So compared with the other two methods, single-line straight-through method, which has good stability and high knitting efficiency and is easy to be realized on the machine, is the focus of the discussion below. Based on this method, an automatic weaving device capable of weaving a beaded cool pad is proposed and designed. 4.1 Feeding Device Design Before weaving the beaded cooling pad, since all the beading arrangements are disordered, it is necessary to design a device that puts the beads into the weaving state in an orderly manner, which is called feeding device. Referring to the feeding mechanism of the firecracker weaving [14-16], Fig.14 shows the schematic diagram of the designed bead feeding mechanism device. Before the device runs, all the beads are placed in the hopper and the two guiding wheels. A small number of longitudinally are placed in a horizontal arrangement. When the two guide wheels rotate in opposite directions, the beads in the hopper will be putted into the guide groove and conveyed to the front of the beading device in an orderly manner. In order to avoid a rigid collision between the feeding device and the ball transported device, the end of the guiding groove is made by a material with better elasticity. 4.2 Design and Working Principle of the Beaded Pad Weaving Device Fig. 15 demonstrated the beaded pad weaving device. Since the figure is only for explaining the movement process of the beaded pad weaving device, the feeding device is not shown in the figure. And there are 7 motors in this device. Control motor A controls the movement of the threading device. Control motor B controls the movement of the movable line of the downlink line. Control motor C controls the movement of the movable stroke switch. Control motor D controls the rotation of the output port and the braided port. Linear motor E S. Ouyang et al.2544 drives the up-line feed ball push block movement. Linear motor F pushes the braided beaded cool pad unit into the braided port. Linear motor G drives the linear motion of the downlink line feed bead block. Before the device is operated, the downlink threading is first performed. After the downlink threading is finished, the string is installed into the beaded pad weaving device. The downlink line with heavy beads at the end is wrapped around the fixed pulley mounted on the frame. Then it passes through the downlink line movable seat, the through hole, the downlink end sleeve, the bead and the braided port successively, to reach the uplink line. And the uplink line is directly connected to the needle of the threading device, the first string of beads are moved to the corresponding position on the weaving port, as shown in Fig.16. After the string installation is completed, the motor A is manually controlled to make the driving roller be located between the two trapezoidal blocks on the movable seat rail of the threading device. The manually controlled linear motor E is to drive the uplink line to send th e beads push block moves, which is external bead conveyed from the feeding device. It causes the holes axis of the bead to coincide with the needle axis of the bead threading device and the up-line bead push block is in the beading state. The specific work ing process of the device is as follows: Method Research and Mechanism Design of Automatic Weaving\u2026 2545 Step 1: Under the drive of the control motor A, the bead threading device moves to the right. When the driving roller moves in the second trapezoidal block on the movable seat rail, the location clamping position of the needle will be changed. As a result, the uplink line smoothly penetrates an external bead provided by the uplink line bead transported device, as shown in Fig.17. Step 2: The bead threading device continues to move to the right. When the movable seat contacts the movable travel switch, the uplink line is just tightened. At this time, some of the motor operation will change as follows: Control motor A reversed means the bead threading device starts to move to the left. Controlling motor B rotated forward means the downlink line movable seat moves to the right for a suitable distance, providing two beads re quired for the next unit downlink line weaving. After that, controlling motor B stops. Linear motor F runs, the weaving beads are pushed into the weaving port and the output port, then moves back to the initial position. Linear motor E reversely drives, the pushing block of uplink line feeding bead returns to the initial position, and is on out feeding condition, as shown in Fig.18. S. Ouyang et al.2546 Step 3: Under the driving of the control motor A, the bead threading device moves to the left. When the threading device contacts the fixed stroke switch, some of the motor operation will change as follows: Linear motor G forward drives, the pushing block of the downlink line feeding bead pushes a shared bead to make the hole axis of the shared bead coincide with the needle axis. Control motor A rotated forward means the bead threading device starts to move to the right. Linear motor E drives forward, the push block of the uplink line feeding bead is in the feeding state. Control motor C rotated forward, the movable travel switch moves to the left for a suitable distance exactly equal to the length of the string required to weave every bead pad unit, as illustrated in Fig. 19. Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state." + ] + }, + { + "image_filename": "designv11_80_0000813_iemdc.2019.8785258-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000813_iemdc.2019.8785258-Figure4-1.png", + "caption": "Fig. 4. PWM triangle carrier waveform of the first and second windings of a dual three-phase motor with carrier phase-shift angle \u03b3.", + "texts": [ + " The applied voltage must be calculated to analyze the mutual effect between the first and second three-phase windings. We used the following method to obtain the PWM phase voltage waveform, except for the controller and inverter properties, by FEA. We first calculated the current references Id and Iq at the target operation point. Next, we applied the sinusoidal current references that are transformed from the Id and Iq to FEA calculation. We then compared the resulting output phase voltages (Fig. 3) and the triangle carrier wave (Fig. 4) to create the PWM voltage waveforms shown in Fig. 5. Finally, we applied the obtained PWM voltage to FEA again to calculate the carrier harmonic current and carrier harmonic flux density. Note that, since this motor has dual three-phase windings, each winding has a \u03c0/6-rad phase difference in the fundamental voltage reference. However, we can choose the phase-shift angle between each triangle waveform of each voltage reference. Fig. 1. Ratio of carrier harmonic phase voltage amplitude to DC voltage versus modulation depth for the first- to the fourth-carrier harmonics\u2019 nearest sideband" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure1-1.png", + "caption": "Figure 1. The geometrical dimensions of impeller.", + "texts": [ + " It is characters for the ability that producing castings with smooth surface and high accuracy of expected size[1.3]. Moreover, its excellent adaptability to complex sharp of patterns makes the investment casting becoming the preference for alloy impeller fabrication[4]. Furthermore, as the wax mode manufacture is proofed for its longer time consuming when making assembly mode[2], 3D printing for producing patterns attracts many attention for discussing the probabilities of it[3.5.9]. The detail size of the impeller is shown as Figure 1. Alloy ZL104 is planned as the cast material and PLA is used for 3D printing pattern. Outside diameter of the impeller is 60mm while height of center ring is 8.7mm. Meanwhile, thickness of the blade is decided as 2mm according to the Table 1[1]. The distorting blade will leads to fault casting in forms of shrinkage defects and gas porosity[8]. Thus setting the reasonable process by simulation is necessary. ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003964_j.ijleo.2020.165806-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003964_j.ijleo.2020.165806-Figure1-1.png", + "caption": "Fig. 1. diagram of the ODM system.", + "texts": [ + " With a simplest case, the fitting of spherical surface as an example, at least 5 points are required to fit the coordinates of a reference point, and at least 3 points are required to solve the pose, that is to say, for each end pose, at least 15 points need to be measured to solve its corresponding actual pose, obviously, this method is both time and energy consuming. Inspired by the laser 6-dimensional measurement system [7,8], this study attempts to construct a contact-type orthogonal displacement measuring system (hereinafter referred to as the ODM system) to carry out the measurement and calculation of the pose of moving platform of parallel mechanism. Fig. 1 shows a schematic diagram of the ODM system, which includes the reference block to be measured (namely the target reference block), and the displacement sensors, etc. Wherein the target reference block is fixed on the moving platform, several displacement sensors are placed in the three orthogonal directions of the target reference block, all displacement sensors are fixed on the sensor stand, and there\u2019s no relative movement between the sensor stand and the fixed platform; under the action of the elastic force, the displacement sensor is always in contact with the target reference block" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002774_0954406220925843-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002774_0954406220925843-Figure5-1.png", + "caption": "Figure 5. Link coordinate systems for the ROKAE serial manipulator.", + "texts": [ + " It is proved that the multilevel optimization criterion proposed in this paper is effective for the KUKA LBR iiwa serial manipulator and that it helps resolve the problem of measurement configuration optimization. (2) Calibration experiment of ROKAE manipulator Considering the influence of non-geometric error sources (such as joint clearance, thermal expansion, etc.), which are difficult to model correctly and completely in numerical simulation of the positioning accuracy of the serial manipulators, the calibration experiment was physically carried out with a 6-DOF ROKAE serial manipulator. Each link coordinate system was established as in Figure 5, and the nominal MD\u2013H parameters were as shown in Table 7. The experimental system and measurement process are shown in Figure 6. Spherically mounted reflector was installed at the end of the serial manipulator flange. The end position of the ROKAE serial manipulator was accurately measured by a Leica laser tracker. The positioning error of the serial manipulator was calculated by the difference between the measured end position and the nominal position obtained by the controller. The measurement configurations optimized by indexO1 were adopted as control groups, and the corresponding evaluation indexes of the optimal set and O1 set were as shown in Table 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003517_1464419320955114-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003517_1464419320955114-Figure1-1.png", + "caption": "Figure 1. Electric vehicle powertrain driveline and photograph of 60 kW motor along with drive end bearing.", + "texts": [ + " In this paper, mathematical equations are derived based on the bearing dynamics presented in Upadhyay27 and Harris.28 The objective of the study is to understand the impact on the acoustic behaviour of bearing. The proposed model can evaluate the fault frequency and an acoustic level. In the present study, validation of a mathematical model has performed with the help of systematic experiment. An adequate correlation between theoretical and experimental results has presented in this research work. Schematic Diagram of Induction motor of 60 kW capacity is shown in Figure 1. The deep groove ball bearing is mounted on a drive and non-drive end position. The major focus of this study is Non-drive end side bearing, which creates huge disturbance if pre-dispatch handling damage occurs. Free body diagram with coordinate system X and Y direction are shown in Figure 2(b) and (c). During design of the mathematical model, the following realistic assump- tions are considered:- \u2022 The housing, bearing and shaft are modelled using three DOF system. \u2022 Forces due to Centrifugal phenomenon on the balls are negligible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000967_iccar.2019.8813340-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000967_iccar.2019.8813340-Figure1-1.png", + "caption": "Figure 1. Installing a bearing loader near outboard bearing housing.", + "texts": [ + " Firstly, the vibrational data were recorded with no load applied on the rotor and the healthy bearings were installed at outboard and inboard bearing housings, and then a healthy bearing at outboard bearing housing was replaced with each faulty bearing. Secondly, a constant load is applied by placing a bearing loader at a distance of 5 cm from the outboard board bearing housing and the vibrations were recorded for the healthy bearings, then again the outboard\u2019s bearing is replaced with each faulty bearing. The installation of the bearing loader is illustrated in Figure 1. The faulty bearings (ball, outer, inner, combination) were only installed at outboard bearing housing and octave plots of faulty bearings have been compared with a healthy bearing at each rotational speed. All of the vibrational data were recorded through the accelerometers mounted at the inboard and outboard bearing housing in horizontal and vertical directions. These vibrational data have been acquired with 10 kHz maximum frequency and 12800 spectral lines. The ISO standard 20816-1 [21] has been followed in this paper by calculating the overall velocity-RMS at each rotational speed and with both types of loading conditions before creating any fault" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003657_j.matpr.2020.08.023-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003657_j.matpr.2020.08.023-Figure3-1.png", + "caption": "Fig. 3. Manufactured links: volu", + "texts": [ + " However, the performance values of the manipulator-link are within acceptable limits. The obtained topology from MATLAB result is imported to SOLIDWORKS software to convert topology from 2D to 3D at the desired thickness. Computer numerical control milling machine is used to manufacture the optimum link with the help of G codes developed in MASTERCAM software. Based on the stated process, the manipulator-links are manufactured for volume fraction 1.0 (solid link) and 0.5 (optimum link), as shown in Fig. 3. To validate the simulation results and to find out the current supplied to the servo motor, an industrial level automation system is developed. The PLC is employed to create the logic for controlling the servo motor through the ladder diagram. HMI is used for ease Table 1 Performance values of the manipulator-link. Volume fraction Deflection (mm) Stress (MPa) 1.0 0.00398 3.62 0.5 0.00983 6.29 operation of the tags provided by the PLC from the human end. The control units are communicated through PROFINET, which is commonly connected to the computer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002406_icuas.2017.7991404-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002406_icuas.2017.7991404-Figure1-1.png", + "caption": "Fig. 1: Quad-rotor helicopter scheme.", + "texts": [ + " Section VI presents the application of the composite control law to the quadrotor multi-agent system. Section VII presents the experimental platform used in this work and the real-time results. Finally, some conclusions and outlooks are offered in Section VIII. 978-1-5090-4494-8/17/$31.00 \u00a92017 IEEE 1324 Consider a quad-rotor helicopter moving in a three dimensional space. Consider now an North-East-Down (NED) inertial frame fixed to earth Fe and a body frame fixed to the quad-rotor helicopter Fb, as shown in Fig. 1. The position of the center of mass of the quad-rotor helicopter with respect to Fe is given by \u03be = [x, y, z] T and its orientation is given by \u03b7 = [\u03c8, \u03b8, \u03c6] T , where \u03c8, \u03b8, and \u03c6 are the Euler angles of yaw, pitch and roll, respectively. Based on [10], the nonlinear dynamics of the quad-rotor helicopter can be expressed as m\u03be\u0308 =\u2212mgD + RF (1a) I\u2126\u0307 =\u2212 \u2126\u00d7 I\u2126 + \u03c4 (1b) where D = [0, 0,\u22121] T , R \u2208 SO(3) is a rotation matrix relating Fe with Fb, F = [0, 0, u] T represents the vector of forces applied to the quad-rotor helicopter, with u =\u22114 i=1 Fi being the main thrust" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002638_pesgre45664.2020.9070716-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002638_pesgre45664.2020.9070716-Figure8-1.png", + "caption": "Fig. 8. Flux density for 40 W (16 Pole) SPIM under locked rotor condition", + "texts": [ + " This power from test results is higher than the experimental difference in power. The core losses under the locked rotor condition are 2.5386 W, as estimated from FEA. However, the core losses of the motor under the no load condition are 5.7695 W. The core is highly saturated under the no load condition as, significant magnetization current is drawn. Under locked rotor condition, the rotor resistance is comparable to the stator resistance. Hence, the current drawn from supply majorly heats up the stator (both the windings) and rotor resistance. Fig. 8 shows the flux density plot for 40 W SPIM motor under locked rotor condition. The stator and rotor core are unsaturated as seen in the plot confirming with 2.5386 W core losses. The rest of the losses are the copper losses (57.10 -2.5386 W= 54.5614 W). Fig. 9 shows the flux density plot of the 40 W SPIM motor under no load condition. The stator as well as rotor core are highly saturated confirming higher losses. The performance of the fan at desired speed (for a particular blade) is calculated from FEA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002114_icar46387.2019.8981663-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002114_icar46387.2019.8981663-Figure2-1.png", + "caption": "Fig. 2: The overall hardware setup used for the interaction between HoLLiE and the user.", + "texts": [ + " The feedback about the robot behavior and the tracking of the tools enable him to feel more comfortable during the control and achieve a better precision. For the interaction control developed we used our service robot HoLLiE, which has two PILZ PRBT robotic arms with 6-DOF. For the interaction system developed, we focused on the control of one arm, which is equipped with a Schunk SVH 5-finger hand. The setup included the marker based tracking system OptiTrack and a monitor to display the user interface. The overall hardware setup is depicted in Fig. 2. Fig. 3 represents the system software architecture and the communication between the components. We used the behavior framework FlexBE [24] to define the high level logic of the interaction and the handling of the events coming from the other modules. The control of the robot arm is done by alternating between a joint-position controller and the Cartesian controller described in [25], implemented with ROS control [26]. The joint-position controller is used to autonomously reach, grasp and dispose tools with fixed trajectories defined with our Motion Pipeline" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001406_012064-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001406_012064-Figure1-1.png", + "caption": "Figure 1. The design of a three-electrode plasma torch with alternating current.", + "texts": [ + " Combining the functions of plasma surface cleaning and laser hardening of metals leads to an increase in the efficiency of the TP. To evaluate the results of plasma cleaning, a series of experimental studies on the deposition of BoroTec - Eutalloy\u00ae 10009 powder on a worm milling cutter made of steel R18K5F2 was carried out. Based on the conducted patent information review on plasma technologies and TP purification, a three-electrode AC plasma torch was developed (Pat. No. 2558713, a device for a pulsed plasma alternating current generator) (Fig. 1). An experimental installation of a PTS based on a threeelectrode alternating current plasma torch was developed, which allows cleaning the surface of metals in a short period of time, which leads to an increase in the efficiency of the TP [2-4]. The performance characteristics of parts are largely determined by the state of the surface and its physicochemical properties. Studies on the treatment of the metal surface with a pulsed plasma flow show a change in the purity of the surface layer, which leads to an improvement in the quality indicators of heat-strengthened surfaces [4]", + " The deviation of the PTS parameters from the optimal values contributes to the intense erosion of both the main electrodes and the nozzle insert, which leads to the destruction of the structural elements of the plasma torch. Technological characteristics of the plasma cleaning process in pulsed modes depend on the energy parameters of the plasma torch, pulse duration, processing schemes, type of material being processed and other factors [8-9]. The design of a three-electrode plasma torch with alternating current is shown in Figure 1. The design (Fig. 1) includes a preliminary gas treatment unit 1 with a screw nozzle of variable section 8, a discharge chamber 2 in which three electrodes 3 are fixed, a water-cooled nozzle 4, an auxiliary pulse source of arc start 5, ceramic bushings insulators 6, fittings 7 for supplying cooling water and plasma gas. The electrodes of the plasma torch 3 form a funnel-shaped three-beam shape and are arranged at an angle of 2 \u00b0 to the axis of the plasma torch. During the development of PTS, studies were conducted to identify the dependence of the quality indicators of cleaning and heat treatment of the surface of parts on the flow of nitrogen in the plasma torch" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003742_icarm49381.2020.9195386-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003742_icarm49381.2020.9195386-Figure3-1.png", + "caption": "Fig. 3. Kinematic diagram of the orthosis coupled with the ankle foot complex.", + "texts": [ + " With the help of the spherical joints, mainly joints S4 and S7, the axial orientation of the device is able to change gradually and finally aligns with that of the user; the dashed and solid figures are the initial and final positions of the device respectively. Therefore, the device has the potential to adapt to users with different orientations of joint axes. III. VELOCITY TRANSMISSION RATIO In order to make the description of this section clearer, the kinematic diagram of the AFO coupled with the ankle and foot complex with only one position is shown in Fig. 3. The diagram is divided into two closed motion chains. Specifically, chain 1 starts from Base, Link 1, Link 2, Link 3 , Link 10 (front foot), Link 9 (talus) and Base (shank); Chain 2 is mainly composed of Base, Link 7, Link 5, Link 4, Link 6 and Link 8. Since the motion of ankle and foot complex is mainly affected by chain 1, in the following motion analysis, chain 1 is focused on. Actually, the objective of motion analysis is to calculate velocity transmission ratio between Link 1 and Link 9. During normal walking, the rotation motion mainly occurs at the ankle joint. Therefore, in this analysis, in order to simplify the derivation, the motion of subtalar joint is neglected. Based on above simplifications, the kinematic diagram of the motion system can be described as Fig. 4. The system has four links, Link 1\u2019, 2\u2019, 3\u2019 and 4\u2019; they are equivalent to Base, Link 1, Link 2 and Link 9 in Fig. 3 respectively. Additionally, the system consists of two rotational joints and two spherical joints; they are R1, R2, S1 and S2 respecitvely. Accordingly, four coordinate systems are defined. Specifically, as shown in Fig. 4, o1, o2, o3 and o4 are the original points of the 978-1-7281-6479-3/20/$31.00 \u00a92020 IEEE 583 Authorized licensed use limited to: University of New South Wales. Downloaded on November 15,2020 at 09:08:52 UTC from IEEE Xplore. Restrictions apply. four coordinate systems respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001199_1.j057603-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001199_1.j057603-Figure5-1.png", + "caption": "Fig. 5 Disk rolling without slipping on a horizontal plane.", + "texts": [ + " (24) yields the following expression of the canonical motion variables column matrix w: w Q 1\u22152 h u m2 p u1 m2 p u2 m1 p u3 T (102) Therefore, the kinematical subsystem in terms of the canonical motion variables is obtained as _q Cu Cww (103) where Cw CQ \u22121\u22152 h 2 666666664 0 0 \u2212 1 L1 m1 p 0 \u2212 1 L2 m2 p sin q2 sin q1 L2 m1 p sin q2 1 L3 m2 p cos q3 0 0 3 777777775 (104) The constraint matrices A1 and A2 are given by A1 A1Q \u22121\u22152 h 2 4\u2212 1 m2 p A 1;1 1 m2 p 0 \u2212 1 m2 p A 2;1 0 1 m1 p 3 5 (105) and A2 A2Q 1\u22152 h m2 p m2 p A 1;1 m1 p A 2;1 (106) and the MPGIs A 1 and A 2 are given by A 1 AT 1 A1AT 1 \u22121 1 \u03b7 2 666666664 \u2212 1 m2 p A 1;1 \u2212 1 m2 p A 2;1 1 m2 p 0 0 1 m1 p 3 777777775 \u00d7 2 6664 1 m1 1 m2 A2 2;1 \u2212 1 m2 A 1;1 A 2;1 \u2212 1 m2 A 1;1 A 2;1 1 m2 1 m2 A2 1;1 3 7775 (107) \u03b7 1 m2 1 m2 A2 1;1 1 m1 1 m2 A2 2;1 \u2212 1 m2 2 A2 1;1 A 2 2;1 (108) and A 2 AT 2 A2AT 2 \u22121 AT 2 A2AT 2 1 m2 m2A 2 1;1 m1A 2 2;1 2 4 m2 p m2 p A 1;1 m1 p A 2;1 3 5 (109) The CEOMs are given by Eq. (43), where B is given by Eq. (82), and D ow nl oa de d by U N IV E R SI T Y O F C A M B R ID G E o n O ct ob er 2 3, 2 01 9 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .J 05 76 03 V Q \u22121\u22152 h F \u2212 L 0 m2 p g m1 p g sinq1 T (110) and u Q \u22121\u22152 h w w1 m2 p w2 m2 p w3 m1 p T ; q2 \u2260 0; q3 \u2260 \u03c0\u22152 (111) The disk shown in Fig. 5 is rolling on a horizontal inertial plane. It is assumed that the center of gravity of the disk coincides with its geometric center point C, and that the friction between the disk and the plane is sufficiently big such that the disk is rolling without slipping on the plane. It is also assumed that the disk is extremely thin such that, at every instant of rolling time, it contacts the plane at a single point B of the circular edge of the disk. The orientation of the disk frameDwith respect to the inertial frameE is taken to followa 3- 2-1 Euler\u2019s sequence of rotations that is made by first translating the origin O of E to B and rotating E about its ZE axis by a precession angle \u03c8 , arriving at the auxiliary reference frame T and then rotating T about its YT axis by a nutation angle \u03b8, arriving at the auxiliary reference frame S, and finally translating the origin B of S to C and rotating S about its XS axis by a rolling angle \u03d5, arriving at the physical reference frame D", + " (117) for v1, v2, and v3 yields u4 cos q2 cos q3 _q4 cos q2 sin q3 _q5 \u2212 sin q2 _q6 (146) u5 \u2212 cos q1 sin q3 sinq1 sin q2 cos q3 _q4 cos q1 cos q3 sinq1 sin q2 sin q3 _q5 sin q1 cos q2 _q6 (147) u6 sin q1 sinq3 cos q1 sin q2 cos q3 _q4 \u2212 sin q1 cos q3 cos q1 sin q2 sin q3 _q5 cos q1 cos q2 _q6 (148) Substituting the preceding written expressions of u4, u5, and u6 in Eq. (145) yields _q6 \u2212 sinq2r u2 cos q1 \u2212 u3 sin q1 0 (149) which can further be reduced via Eq. (116) to _q6 \u2212 r sin q2 _q2 0 (150) The preceding constraint equation is time integrable to q6 r cos q2 c (151) where c 0 because q6 at q2 0 equals to \u2212r, as depicted from Fig. 5. Therefore, the second constraint in Eq. (143) is holonomic, and it is the constraint that accompanies the involvement of q6 as the pseudogeneralized coordinate among the six generalized coordinates. The AFC of this second constraint is obtained by taking the time derivative of Eq. (149) as _u6 \u2212 r _u2 sin q2 \u2212 ru22 cos q2 0 (152) To impose the no-slip condition EvB 03 on the true (threeDOFs) disk, it remains to satisfy the physical constraint EvBf \u22c5 t2 0 (153) on the motion of point B, i.e., u5 cos q1 \u2212 u6 sinq1 \u2212 ru1 0 (154) which is a nonholonomic constraint because it cannot be put in an algebraic closed form that involves generalized coordinates but no motion variables" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001324_chicc.2019.8866109-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001324_chicc.2019.8866109-Figure1-1.png", + "caption": "Fig. 1. Quadrotor configuration frame scheme with body fixed and the inertia frames.", + "texts": [ + " In section 2, the mathematic model of an X-type quadrator UAV is established. Section 3 proposes a cascade PID control and an ADRC strategies for quadrotor UAV. Simulation examples are given in Section 4 to demonstrate the effectiveness of the proposed technique, and Section 5 concludes the paper. In this section, we give the mathematic model of an Xtype quadrator UAV. The nonlinear dynamics are obtained in East-North-Up (ENU) inertial and body-fixed coordinates. Let us consider earth fixed frame E and body fixed frame B, see Fig. 1. Define \u0398 = [\u03c6, \u03b8, \u03d5]T and \u03c9 = [p, q, r]T , where \u03c6, \u03b8 and \u03d5 denote the angle of roll, pitch, and yaw with respect to the inertia frame and p, q, r denote the angular velocity of roll, pitch, and yaw with respect to the body-fixed frame. The rotation matrix from body-fixed frame to inertia frame can be obtained as Rt = \u23a1 \u23a3 C\u03d5C\u03b8 C\u03d5S\u03b8S\u03c6 \u2212 S\u03d5C\u03c6 C\u03c8S\u03b8C\u03c6 + S\u03c8S\u03c6 S\u03d5C\u03b8 S\u03d5S\u03b8S\u03c6 + C\u03d5C\u03c6 S\u03c8S\u03b8C\u03c6 \u2212 C\u03c8S\u03c6 \u2212S\u03b8 C\u03b8S\u03c6 C\u03b8C\u03c6 \u23a4 \u23a6 , (1) where S(\u00b7), C(\u00b7) and T(\u00b7) denote sin(\u00b7), cos(\u00b7) and tan(\u00b7), respectively. According to the rotation matrix Rt, the relationship between \u0398\u0307 and \u03c9 can be described as \u23a7\u23a8 \u23a9 \u03c6\u0307 = p+ T\u03b8S\u03c6q + T\u03b8C\u03c6r \u03b8\u0307 = C\u03c6q \u2212 S\u03c6r \u03d5\u0307 = S\u03c6/C\u03b8q + C\u03c6/C\u03b8r" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003423_iccsse50399.2020.9171958-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003423_iccsse50399.2020.9171958-Figure2-1.png", + "caption": "Figure 2. Centroid radius vector of bridge.", + "texts": [ + " Through the cooperative operation of the crane traveling mechanism and the trolley, the lifting equipment covers all areas in the containment vessel. Based on the principle of virtual power, a multi-rigid body dynamic model of polar crane is established. According to a single rigid body dynamic equation: 1 1 1 1 0 0 a c a m r m g F J J M (1) In the Eq. (1), aF is the external force on the rigid body, aM is the moment of the external force on the rigid body around the center of mass, 1J is the moment of inertia of the rigid body. Establish the vector relationship as shown in the fig. 2. The virtual work equation of each object in the polar crane can be expressed as: 0 0= T i i i r r P M F (2) where 0r is the vector from global coordinate to local coordinate of bridge, cr is the vector from global coordinate to the bridge center of mass, icr is the vector of each object center of mass in the bridge local coordinate, icr is the antisymmetric matrix of icr , E is the unit matrix, and = ic ic ic ic mE mr M mr J mr r (3) = ic ic ic ic mg m r F mr g J mr r (4) According to the dynamic modeling process of literature [7], the overall dynamic model of the polar crane can be obtained as follows: Mq F Kf (5) where 31 1 = T i i i i M T M T , 31 1 = T i i i i i F T F M W are the generalized mass matrix and generalized force matrix of the system, 1,2 31i is the number of each object in the polar crane, q is the acceleration vector for generalized coordinate, Kf is the wheel-rail generalized contact force transformation matrix, is the wheel-rail contact force matrix, iT is the generalized coordinate transformation matrix of each object, and iW is the remainder term of acceleration vector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000279_j.procs.2019.02.074-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000279_j.procs.2019.02.074-Figure3-1.png", + "caption": "Fig. 3. (a) \u2013 the kinematic scheme of the manipulator; (b) \u2013 structure of the TR control circuits.", + "texts": [ + " Consider the implementation of the correction contour of the perturbations due to the unaccounted technological factors 3 1 1 1( , ..., , \u03c9 , ..., \u03c9 , \u03c6 , ..., \u03c6 )n n nA AF using the example of a multi-axis technological manipulator. This technological robot has six degrees of mobility 1 6( , ..., ),q q i. e. while the generalized coordinates 1 2 3, ,q q q form the carrier system of the mechanism, and 4 5 6, ,q q q are responsible for the orientation of the TR working body in space. The kinematic scheme of this TR is shown in Fig. 3a. For simplicity, we will consider the interaction of the actuators responsible for the angular displacements along the generalized coordinates 2 3,q q which corresponds to the contours 2, 3 (Fig. 3, b). Moreover, when they are cross-adjusted, we will consider the dynamics and accuracy of the TP as a whole. In the simulation, a constant perturbation was given in the form of an additional component for the output coordinate of the second link. Using the actuator model by the perturbation 2 ( )M p this component 2f is extracted in the second loop. Further through the correction circuit 23 ( )K p is introduced into the third circuit, as an additional component 3.g As the models of the contours by the perturbation, we use the static dependencies of the first and second error coefficients on the momentum (the torque)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001373_icmsao.2019.8880400-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001373_icmsao.2019.8880400-Figure1-1.png", + "caption": "Fig. 1. 3DOF laboratory helicopter model [1]", + "texts": [ + " INTRODUCTION The flight control system design has picked up a developing enthusiasm for late years particularly for the 3DOF helicopter system, which is an effective benchmark unmanned aerial vehicle (UAV) model for theoretical and experimental study on flight control algorithm. Large levels of maneuverability and the capacity of working in antagonistic climate conditions are the new patterns in helicopter design these days. Helicopter flight control system should make these execution necessities reachable by enhancing tracking execution and disturbance rejection ability. The helicopter model utilized in this paper is a laboratoryscale 3 DOF helicopter (see Fig. 1) and it is produced by Googol Technology Ltd., Kowloon, Hon Kong. The helicopter model contains two DC motors which drive its two propellers respectively. These motors are mounted at the ends of a rectangular edge. The frame of the helicopter is suspended at one end of a long arm and its other end of arm carries a balance block. The 3 DOF helicopter model contains open-loop unstable dynamics. The desired elevation angle of the helicopter can be obtained through by providing positive voltage to either motor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002873_icpes47639.2019.9105570-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002873_icpes47639.2019.9105570-Figure3-1.png", + "caption": "Fig. 3: Line current density model with excitation A(\u03b3, t).", + "texts": [ + " The distributed winding models are operated with a stator frequency of f = 300 Hz and the fractional-slot winding models with f = 200 Hz. All models are excited with a quadrature current component Iq (Id = 0 A). In [5] a similar model was used to identify the rotor losses of a PMSM. In this analysis the model is used to identify the influence of the different harmonics and sub-harmonics on the iron and eddy-current losses of the stator, rotor and PMs. The model couples the analytical description of the space harmonics in the air gap produced by the stator windings with a 2D-FEA model of a machine without windings and slots. Fig. 3a shows the original machine model and Fig. 3b the derived line current density model. The derived model consists only of an iron ring which represents the stator yoke Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 14,2020 at 15:36:02 UTC from IEEE Xplore. Restrictions apply. without any slots or windings. For the excitation of the model, a line current density A(\u03b3, t) = 2 DS \u2202\u0398(\u03b3, t) \u2202\u03b3 = \u2211 \u03bd A\u0302\u03bd sin(\u03bdp\u03b3 \u2212 \u03c9t) A\u0302\u03bd = \u22122 \u221a 2 I \u03c0DS \u00b7 p \u00b7m \u00b7N1 \u00b7 \u03b6\u03bd , (9) is used, which can be calculated from the spatial distribution of the MMF mentioned in (1)", + " First, the effect of the single harmonic orders is considered. Secondly, the amplitude reduction of the MMF harmonics is shown. In the end, the impact of the subharmonics is explained. A. Influence of the harmonics 1) Fundamental and Harmonic Waves: In order to evaluate the impact of the fundamental wave and the single harmonics on the losses in the stator, rotor and PMs, a non-distributed integer-slot winding (q = 1) is considered. In order to separate the losses into their excitation order, the line current density model from Fig. 3b is used. The MMF spectrum of this winding scheme is shown in Fig. 6. It can be seen that just a high fundamental wave and smaller harmonic orders (\u03bd = (2mg + 1) , g = \u00b11,\u00b12,\u00b13, . . . ) exist. The losses of each single harmonic are visualised in Fig. 7. The first column represents the total losses of the simulation which considered all harmonics. The column is divided into stator, rotor and PM losses. The second column represents the sum of the losses of the single harmonic orders. The losses of the single harmonic orders are presented in the following columns" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002052_b978-0-12-812750-6.00001-9-Figure1.33-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002052_b978-0-12-812750-6.00001-9-Figure1.33-1.png", + "caption": "FIGURE 1.33 Different extent of underride in car-to-truck impact. From Allen, K., 2010. The effectiveness of underride guards for heavy trailers. In: NHTSA Technical Report DOT HS 811 375 (Allen, 2010).", + "texts": [ + " This occurs also in a rear-end impacts, when, for example, the vehicle is braking, see Fig. 1.30. In this case, the damaged area and the acceleration curves are different for engagement or underride impact, as shown in Figs. 1.31 and 1.32. Other underride situations are when a passenger vehicle collides with the rear end or side of trucks, trailers, or buses not equipped with effective guards. In these cases the vehicle continues to travel beneath the taller chassis of the larger vehicle, as depicted in Fig. 1.33. In the case of underride/override the contact will have high losses of kinetic energy due to sliding friction of the sheets between them, even in case of low permanent deformations. This aspect is more evident in lowspeed impact, in which the magnitude of the friction phenomena has a higher relative weight on the entire portion of kinetic energy loss. In the case of underride/override the force/deformation curves are not comparable with the ones obtained in the standard crash test. The F(x) curve, in the typical range of speed that occurs in collisions between vehicles, is nearly independent of the deformation speed, that is, from dx/dt" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002817_tii.2020.2998165-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002817_tii.2020.2998165-Figure3-1.png", + "caption": "Fig. 3. The graphical explanation of the relative insertion depth.", + "texts": [ + " Therefore, the state transition process is modeled as f(k+1) = f(k)+g(\u2206c(k))+ \u03b5\u03b4 (k), (7) where g(\u00b7) is the function describing the effect of relative movements upon interaction forces, \u2206c(k) is the compensation movement increment, \u03b5\u03b4 (k) is an uncertain item relating with movements, and \u03b4 (k) is the one-step relative insertion depth vector. Given the relative movements of all objects, this equation approximates how interaction forces change. The stochastic element \u03b5\u03b4 (k) is a Gaussian vector whose parameters depend on the insertion depths between pegs and holes. For the movement x(k), the relative insertion depth results in \u03b4i(k) = [(\u2206xi(k)\u2212\u2206xi+1(k)) \u00b7hi]hi, (8) where \u03b4i is a subvector of the vector \u03b4 , \u2206xi(k) is the object movement increment, and \u201c\u00b7\u201d is the dot product. The graphical explanation of \u03b4i is shown in Fig. 3. The uncertain item then conforms to \u03b5\u03b4 (k) \u223c N ( \u00b5\u03b4 (k),\u03a3\u03b4 (k) ) , (9) where \u00b5\u03b4 (k) and \u03a3\u03b4 (k) are the mean and the covariance corresponding to the relative insertion depth. It shows that the Gaussian parameters in the state transition process vary with movements, and this varying rule should be obtained via experiments in advance [8]. The effect of relative movements upon interaction forces is an important aspect and, to approximate it, we separate it into two parts g(\u2206c(k)) = gr (\u2206c(k))+ga (\u2206c(k)) , (10) where gr(\u00b7) = [gT r1(\u00b7),gT r2(\u00b7), \u00b7 \u00b7 \u00b7 ,gT rn(\u00b7)]T and ga(\u00b7) are two functions relating to radial and axial forces with respect to their own peg directions, individually" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure9.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure9.10-1.png", + "caption": "Fig. 9.10 FE simulation results of residual stress by DCR to 0.5 mm tool depth", + "texts": [ + " 9 Experimental and Numerical Assessment of Residual Stresses \u2026 105 This could be attributed to the higher plastic deformation resulting in higher dislocation density at a distance of 0.15 mm below the surface, and the material is getting plasticised between the tool and the material itself due to resistance to deformation from the strength offered by the material below the subsurface. A similar trend in the results of residual stresses is observed with different tool penetration depths by DCR of 0.5, 1.0, 1.5 and 2.0 mm. The contour of longitudinal residual stress obtained from simulations is shown in Fig. 9.10. It is found that the stresses on the surface of the specimens in the un-deformed region are tensile in nature whereas the stresses on the deformed region are compressive. This could be attributed to self-equilibrating nature of residual stress, i.e. the net sum of residual stress is zero. 106 R. Kumar et al. Aluminium plate of AA6061 with 8 mm thickness is deep-rolled to different depths of 0.5, 1.0, 1.5 and 2.0 mm. On the basis of experimental work and FE simulations, the following conclusions can be drawn: \u2022 The yield strength of aluminium alloy AA6061 in annealed state is 69 MPa and tensile strength is 132 MPa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000149_robio.2018.8664847-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000149_robio.2018.8664847-Figure4-1.png", + "caption": "Fig. 4. An encountered-type haptic interface using MR fluid", + "texts": [ + "5 [N] force at the maximum in the cutting direction occurred. Besides, Figure 3(b) shows that 1 - 3 [Hz] frequencies are mostly contained while more than 10 [Hz] frequencies are almost not observed in the signal. From this result, we aimed to develop the haptic interface using MR fluid which satisfies the following required specification. 1) the maximum generating force is more than 1.5 [N] 2) the resolution of the generating force is less than 0.1 [N] and more than 10 [Hz] frequency response is required. Figure 4(a) shows an overview of the encountered-type haptic interface using MR fluid. This interface consists of a container filled with MR fluid and dual coils. By controlling external magnetic field, apparent viscosity of MR fluid is changed depending on the magnitude of magnetic field. Then, an operator can feel resistance force by cutting the MR fluid directly. To reproduce cutting resistance force of biological tissue, the following force feedback control is applied to this system. I = KP\u0394F +KD\u0394 F dt + I \u2032 (1) where I , Kp and KD are command current of coils, proportional gain and derivative gain, respectively. \u0394F is deviation between reference cutting resistance force and measured cutting force and I \u2032 is an offset current. As shown in Figure 4(b), cutting force is measured by a surgical instrument equipped with force sensors [5]. The measured force is fed back to a microcomputer and reference current is calculated with PD feedback based on eq. (1). Finally, PWM signal is generated to control current to the coils. The above developed system is able to measure the cutting resistance force with 0.1 [N] resolution and to perform approximately 10 [Hz] of frequency response [5]. This indicates that the system has sufficient performance to reproduce vibratory cutting force of biological soft tissue" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000972_00207721.2019.1655601-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000972_00207721.2019.1655601-Figure10-1.png", + "caption": "Figure 10. Simplified schematic diagram of the 1-DOF flexible robot arm control system.", + "texts": [ + "22 3\u2211 i=1 3\u2211 j=i s ,1i,j \u239e\u23a0 = 0. The root z = \u22121.2438 is found by using Tables 1 and 2, meaning the controller designed in Example 1B now becomes unstable. The control performance and controller output are shown in Figures 8 and 9. In Example 2, we apply the proposed control scheme to a one-degree-of-freedom (1-DOF) flexible robot arm control system, which further illustrates our designed controller. Example 2 (Part A): The simplified schematic diagram of the 1-DOF flexible robot arm control system is shown in Figure 10, and its dynamic model is as follows (to Figure 8. Output of the unstable system in Example 1A controlled by the controller designed in Example 1B with rf being 1.2 instead of 0.7 (Figure 5). Figure 9. Output of the controller designed in Example 1B with rf being 1.2 instead of 0.7 (Figure 6). lighten the notation, from now on to the end of Example 2, Example 2A (subsequent Example 2B) will be employed instead of Example 2 Part A (Example 2 Part B)): J \u03d5\u0308 = \u2212MGR cos\u03d5 \u2212 C\u03d5\u0307 \u2212 K\u03d5 + \u03c4 , (26) where \u03d5 is the measured angle of the arm, whose zero position is horizontal and positive direction is counterclockwise, \u03c4 is the control torque, and model parameters are given in Table 3", + " system is y(k + 2)\u2212 \u22112 =1 h ( 0) (\u22112 i=1 a i y(k + 2 \u2212 i) ) = 0, which becomes y(k + 2)\u2212 1.9922y(k + 1)+ 0.9996y(k) = 0 by using Equations (32)\u2013(35). The corresponding z-transform equation is z2 \u2212 1.9922z + 0.9996 = 0. The two roots are z1,2 = 0.9996 \u00b1 0.0200i, which are inside the unit circle, hence the system is stable at y(k + 1) = 0. It can be verified through simulation by assuming a very small initial value for the system output (see Figure 11). The result is consistent with the analysis of the stability of the zero equilibrium point by the physical model in Figure 10, and the quality of the established discrete-time quadratic TS fuzzy model is verified. Example 2 (Part B): Because nu = 0 in Example 2A, we expect to design a feedback linearisation controller for the stable system in Example 2A to achieve perfect tracking of the following constantly-varying differentiable reference trajectory: r(k) = \u2212\u03c0/4 + (1 \u2212 cos(2kT/3)). (36) Solution: Because \u22112 =1 h ( y(k + 1))g 0( y(k + 1)) = T2 J = 0, condition (13) in Theorem 3.3 holds. Consequently, we can utilise the original controller in Theorem 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002982_j.mechmachtheory.2020.104001-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002982_j.mechmachtheory.2020.104001-Figure6-1.png", + "caption": "Fig. 6. SFC with three contacts. (a) Object with three spherical concave contacts, (b) coordinate systems at contacts.", + "texts": [ + " The remaining three non-zero twist coordinates, N , P \u2217a , and Q \u2217 a , makes the pair of objects behave as a 3-d.o.f. planar joint. 5.3. SFC With three contacts Now consider the configuration wherein three n-lines are concurrent and lie in a plane similar to that of Fig. 3 (b). In this example it is shown that an object can be in SFC with three spherical concave contacts. The three points of contact, labeled C a , C b , and C c , are located at the vertices of an equilateral triangle of edge length e . The respective coordinate systems and geometry are shown in Fig. 6 . The second-order motion contact coefficient A for each of the contacts, expressed in terms twist and twist-derivative coordinates in the fixed coordinate systems at respective contacts, are: A a = \u22121 8 (k m + k f ) { L 2 a + M 2 a \u2212 (k m \u2212 k f ) L a Q \u2217 a + (k m \u2212 k f ) M a P \u2217 a \u2212 k m k f P \u22172 a \u2212 k m k f Q \u22172 a \u2212 (k m + k f ) R \u2217 da } (58) A b = \u22121 8 (k m + k f ) { L 2 b + M 2 b \u2212 (k m \u2212 k f ) L b Q \u2217 b + (k m \u2212 k f ) M b P \u2217 b \u2212 k m k f P \u22172 b \u2212 k m k f Q \u22172 b \u2212 (k m + k f ) R \u2217 db } (59) A c = \u22121 (k m + k f ) { L 2 c + M 2 c \u2212 (k m \u2212 k f ) L c Q \u2217 c + (k m \u2212 k f ) M c P \u2217 c \u2212 k m k f P \u22172 c \u2212 k m k f Q \u22172 c \u2212 (k m + k f ) R \u2217 dc } (60) 8 Matrices [ R ab ] , [ R ac ] , [ \u0302 p ab ] , and [ \u0302 p ac ] are obtained as: [ R ab ] = [ cos 120 \u25e6 0 cos 150 \u25e6 0 1 0 cos 30 \u25e6 0 cos 120 \u25e6 ] = [ \u22121 / 2 0 \u2212\u221a 3 / 2 0 1 0 \u221a 3 / 2 0 \u22121 / 2 ] ; (61) [ R ac ] = [ cos 120 \u25e6 0 cos 30 \u25e6 0 1 0 cos 150 \u25e6 0 cos 120 \u25e6 ] = [ \u22121 / 2 0 \u221a 3 / 2 0 1 0 \u2212\u221a 3 / 2 0 \u22121 / 2 ] ; (62) [ \u0302 p ab ] = [ 0 \u2212e \u221a 3 / 2 0 e \u221a 3 / 2 0 \u2212e/ 2 0 e/ 2 0 ] ; [ \u0302 p ac ] = [ 0 \u2212e \u221a 3 / 2 0 e \u221a 3 / 2 0 e/ 2 0 \u2212e/ 2 0 ] (63) The twists $ b \u2261 { L b , M b , N b ; P \u2217 b , Q \u2217 b , R \u2217 b } and $ c \u2261 { L c , M c , N c ; P \u2217c , Q \u2217 c , R \u2217 c } in the coordinate systems at contact- b and contact- c respectively are obtained using adjoint transformation as: [$ b ] = [ L b M b N b P \u2217 b Q \u2217 b R \u2217 b ]T = [( \u2212 L a 2 + \u221a 3 N a 2 ) M a ( \u2212 \u221a 3 L a 2 \u2212 N a 2 )( \u2212 e \u221a 3 M a 2 \u2212 P \u2217a 2 + \u221a 3 R \u2217a 2 ) ( \u2212 e \u221a 3 L a 2 + eN a 2 + Q \u2217 a ) ( \u2212 eM a 2 \u2212 \u221a 3 P \u2217a 2 \u2212 R \u2217a 2 )]T (64) [$ c ] = [ L c M c N c P \u2217 c Q \u2217 c R \u2217c ]T = [( \u2212 L a 2 \u2212 \u221a 3 N a 2 ) M a (\u221a 3 L a 2 \u2212 N a 2 )( \u2212 e \u221a 3 M a 2 \u2212 P \u2217a 2 \u2212 \u221a 3 R \u2217a 2 ) ( \u2212 e \u221a 3 L a 2 \u2212 eN a 2 + Q \u2217 a ) ( eM a 2 + \u221a 3 P \u2217a 2 \u2212 R \u2217a 2 )]T (65) Differentiating the expressions for R \u2217 b and R \u2217c gives: R \u2217 db = \u2212eM da 2 \u2212 \u221a 3 P \u2217 da 2 \u2212 R \u2217 da 2 (66) R \u2217 dc = eM da 2 + \u221a 3 P \u2217 da 2 \u2212 R \u2217 da 2 (67) Analysis for first-order motion space Any allowed first-order motion by the three contacts should satisfy: R \u2217a \u2265 0 , \u2212 eM a 2 \u2212 \u221a 3 P \u2217a 2 \u2212 R \u2217a 2 \u2265 0 , and eM a 2 + \u221a 3 P \u2217a 2 \u2212 R \u2217a 2 \u2265 0 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002087_012059-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002087_012059-Figure1-1.png", + "caption": "Figure 1. Cycloid-pin drive coordinate system", + "texts": [ + " The main factors affecting the transmission accuracy of RV reducer are revealed and the conclusion beneficial to improving the transmission accuracy of ICMAE 2019 IOP Conf. Series: Materials Science and Engineering 751 (2020) 012059 IOP Publishing doi:10.1088/1757-899X/751/1/012059 the system is drawn.In this paper, the geometric characteristics of cycloidal transmission are analyzed to provide guidance for the modification of cycloidal wheel. The principle of cycloidal transmission is shown in Figure.1.The track of the meshing point of the cycloidal pinwheel is a part of the circular arc of the pin teeth in the space coordinate system. Let be the reference coordinate system, is a pinwheel coordinate system, is a cycloid coordinate system.he origin of the cycloid coordinate system coincides with the origin of the reference coordinate system.The distance between the origin of the pinwheel coordinate system and the origin of the cycloid coordinate system is e. is needle roller radius, is the distribution circle radius of pin teeth" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002364_022108-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002364_022108-Figure1-1.png", + "caption": "Fig. 1 Mesh stiffness model", + "texts": [ + " Hereafter, the analytical results are compared with the FEM results in Ref. [11]. Then a 6 DOF gear dynamic is established and the effects of crack depth on the gear dynamics are simulated and investigated. Ease of Use The mesh stiffness of a single helical gear can be defined as the force applied on the contact line over the deflection along the contact line. In order to obtain the mesh stiffness of a single helical gear, a mesh stiffness model is proposed to cut the helical gear into infinite slices along the tooth width as showed in Fig.1. By applying this model, each slice can be considered as a spur gear and the stiffness of each slice can be calculated by energy method [6, 7, and 10]. The deflections of a helical gear slice can be considered as a non-uniform cantilever beam with infinitesimal width dx and effective length d. The bending, shear and axial compressive energy can be expressed by the following equations [6, 7, and 10]. 2 2b b F U k , 2 2s s F U k 2 2a a F U k (1) Where bk , sk , ak represents the bending, shear and axial compressive stiffness in the same direction of the force F", + "1088/1755-1315/440/2/022108 1sinaF F , 1cosbF F (5) 2 2 b 2 2 cos cos d (d r cos ) 0 cos b a b b b a b F x F h r x r M F x F h x r (6) Based on Eqs. (4)- (7), the bending stiffness bk can be calculated by, 2 2 2 cos 1 1 2 20 2 cos 1 b 2 1 20 (cos sin ) cos( ) cos( ) d 1 (cos (d r cos( ) ) sin ) 0 cos( ) b b r d b b x r db b x x h dx r x r EI k x h dx x r EI (7) The shear stiffness sk and axial compressive stiffness ak can be obtained as, 2 2 cos 1 0 1.2cos1 br d xs dx k GA (8) 2 2 cos 1 0 sin1 br d xa dx k EA (9) In the formulas (2)-(8), h, d, x, dx are shown in Fig.1 (b).E denotes the Young modulus. G denotes the shear modulus. xI And xA denote the area moment of inertia and area of the section where the distance between the section and the acting point of the applied force is x, which are showed as follows. 3 3 1 ( ) 12 1 ( ) 12 x hx hx dw hx hc I hc hx dw hx hc (10) ( )d ( )dx hx hx w hx hc A hx hc w hx hc (11) Here, hx denotes the half height of the section where the distance between the section and the acting point of the applied force is x" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003120_s38311-020-0255-4-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003120_s38311-020-0255-4-Figure4-1.png", + "caption": "FIGURE 4 Installation space model of the safety cell (a), result of the topology optimization with combination of all load cases (b), NFRP assemblies of the safety cell (c) (\u00a9 Trier University of Applied Sciences)", + "texts": [ + " The crash elements are designed in a way that the deformation path of the generally larger and heavier accident opponent is initially used for energy dissipation. Through the so-called crushing of tubular or wave-shaped crash elements, a specific energy absorption of approximately 30 kJ/kg is achieved, which is about twice of the specific energy absorption of conventional steel crash boxes [6], as various tests show, FIGURE 3. The development of the monocoque began with the definition of a design space model, FIGURE 4 (a). This was developed on the basis of ergonomic studies of occupants of different statures and optimized in consideration of the drive and chassis package. This was followed by the determination of main load paths using topology optimization, FIGURE 4 (b). This was based on expected loads from the load cases front, side and rear crash and from driving operation. The main load paths were used for the first rough dimensioning of the fiber composite structures and for determining the fiber angles of the unidirectional fibers used, called ampliTex. This structure was then converted into production-ready geometries, FIGURE 4 (c). The proTRon Evolution has a wheel-individual rear wheel drive. This has several advantages: it enables torque vectoring when driving and when recuperating the vehicle, thus ensuring that the vehicle can be kept stable at moderate lateral accelerations even when the rear axle is braked electrically alone. This means that mechanical brakes on the rear wheels can be dispensed with completely and the recuperation potential can be expanded, thus further reducing energy power demand. The drive concept is realized by compact drive set swing arms, which as trailing arms with suspension struts simultaneously take over the wheel guidance and are a novelty in passenger cars, FIGURE 5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003458_med48518.2020.9183024-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003458_med48518.2020.9183024-Figure2-1.png", + "caption": "Fig. 2. Morphing parameter definition", + "texts": [ + " (29) On convergence of the critic network parameter, the actor network parameter is updated using the gradient descent method as follows \u02d9\u0302 WT a = \u2212\u03b1 ( \u03a8(x)T )\u22121 W\u0302T c \u2207u\u03a6T ( x, ( W\u0302 (i) a )T \u03a8(x) ) . (30) Numerical simulation is performed to demonstrate the effectiveness of the proposed control design. To deal with the general nonaffine system, the variable-span and variable-sweep morphing wing aircraft model considering 221 Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 22,2020 at 12:04:42 UTC from IEEE Xplore. Restrictions apply. the morphing parameters as control variables is used. The morphing aircraft model considered in this study is shown in Fig. 2, which has variable-span and variable-sweep morphing parameters. Variable-span and variable sweep morphing are parameterized by two morphing parameters, \u03b71 and \u03b72, where span and sweep angle variations are linearly mapped onto [\u22120.5, 0.5], as summarized in Tables I and II. In this study, the longitudinal motion of the morphing aircraft is considered. The nominal dynamic model is obtained at the flight condition of airspeed 20 m/s with the altitude 300 m, where both morphing parameters are zero. The system model is governed by the following dynamic equations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002392_978-981-15-1293-3_8-Figure6.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002392_978-981-15-1293-3_8-Figure6.10-1.png", + "caption": "Fig. 6.10 (a) An \u2018ABCDE\u2019 coded bio-nano robotic (A, B, C, D and E: constituent assembly bionano modules of the bio-nano robot. (b) Conceptual modular organization result of the bio-nano robot", + "texts": [ + "2), which represent a specific functional module that constitutes the structure of a bio-nano robot. Specifically such as a bio-nano robot ABCD with specific functions defined by each basic bionano module A, B, C, and D.\u00a0For example, a bio-nano robot system with features of energy storage, sensing, signaling and mechanical manipulation at nanoscale, described and represented by code E-S-G-M.\u00a0 Such coding will help to classify, manage, and benefit the mathematical handling of not only the swarms but also bionano robot systems. As in Fig.\u00a06.10b, the illustrated bio-nano robotic system are coded and represented as bio-nano code: EIWR-M-S-FG.\u00a0The conceptual represents modular organization is shown (Fig.\u00a06.10). 6.3 Design and\u00a0Control for\u00a0Bio-Nanorobotic Systems 106 The design, formation, and development of a prototype core universal growth template in not only the bio-nano modularization process, but also the bio-nano robotic system, which could be programmed, upgraded, coded into any bio-nano coded system, and is a notably approach in this area, also called the Universal template or in other words as Bio-nano stem system. It mimics and resembles the concept of stem cells found in the human body. 6 Bio-Nanorobotics: Mimicking Life at\u00a0the\u00a0Nanoscale 107 Internal control mechanisms, which can be achieved through the control of a internal molecular computer, are based on selective biochemical sensing and binding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure6.7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure6.7-1.png", + "caption": "Fig. 6.7 An illustration of aircraft structures using rivets and other fasteners", + "texts": [ + " In the example shown replacing a solid sheet/plate structure of thickness \u2018t\u2019 by a honeycomb sandwich core leads to enhancing the relative bending strength and rigidity several times. Conversely to get the same bending strength and rigidity of the structure as for sandwich construction, the thickness of the solid section will have to be doubled and therefore the weight of the solid structure will be two times higher. Another potential area for light weighting of aircraft is reducing the number of parts. An aircraft structure is an assembly of several component assemblies leading to sub-assemblies and then major assemblies as illustrated in Fig. 6.7. Boeing 787 uses 2.4 million fasteners, about 70% aluminium alloy rivets and 30% high strength steel and titanium alloy fasteners weighing about 12 tons per aircraft at an average weight of 5 g/fastener. A 20% reduction in fasteners required for assembly will reduce weight by 2.4 tons. This can be achieved by making structures by integrally milling from plates instead of riveted assemblies, rivet-less adhesively bonded metal and composite structures, co-cure and co-bonded composites, large single moulded composites by wet-layup, vacuum assisted resin transfer moulding, investment cast moulding for complex pipe assemblies, 3D printing etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001299_aim.2019.8868869-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001299_aim.2019.8868869-Figure13-1.png", + "caption": "Figure 13. Dispensing performance test physical device", + "texts": [ + "93 0 20 40 60 80 100 120 0 50 100 150 200 250 300 \u94f0 \u94fe \u8f93 \u51fa \u4f4d \u79fb \uff08 \u03bc m \uff09 \u9a71\u52a8\u7535\u538b\uff08V\uff09 \u964d\u538b\u66f2\u7ebf \u5347\u538b\u66f2\u7ebf Driving voltage(V) O u tp u t d is p la c e m e n t( m m ) 0 20 40 60 80 100 120 0 5 10 15 20 25 30 35 \u8f93 \u51fa \u4f4d \u79fb \uff08 \u03bc m \uff09 \u9a71\u52a8\u7535\u538b\uff08V\uff09 \u5347 \u964d Driving voltage(V) O u tp u t d is p la c e m e n t ( \u03bc m ) Buck Boost Figure 11. Displacement amplification mechanism hysteresis curve Dispensing jetting test This experiment uses a produced IS-300 series robot, the overall structure of which is shown in Fig. 12. The glue, the flow channel, the striker, and the nozzle are mounted on the three-dimensional motion platform in conjunction with the FHDA, and the glue adopts an industrial resin, and the viscosity at room temperature is 1200 cps. Its overall structure is shown in Fig. 13. Fig. 14 is the piezoelectric ceramic drive voltage signal with the frequency more than 200Hz.The datas is measured by Micsig tablet oscilloscope TO204. The results of the dispensing experiment are shown in Fig. 15. In the figure, the upper and lower rows are respectively effect diagrams for setting the number of dispensing points to 20 and 100 points. We can see that the droplet size is uniform and the amount of glue is the same. Usually the dispensing accuracy is affected by factors such as voltage amplitude, driving air pressure, and nozzle size" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002811_012063-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002811_012063-Figure2-1.png", + "caption": "Figure 2. The assessment set up", + "texts": [ + " The observed system was a prototype as shown in Figure 1. Table was made from metal with size of 500x500x500 mm, two poles with diameter of 20 mm and length of 300 mm and 500 mm were placed on table top. Four units of pillow block NTN UCP 204 were employed with mild steel holders; two spur gears of 33 and 2.5 were attached to the poles. Two 60 mm type-A V-pulleys and a V-belt A39 connects the mechanics to a 3-phase, 1380 rpm, 0.75 kW LM-Motor, which was supplied by 1 HP 3 phase inverter. The bearing assessment set up is shown in Figure 2. Bearings were assessed with speeds and poles axis varied. Vibration was read by using a laser vibration photo sensor, interfaced to a computer by using a labjack that converts analog measurement signals into digital form readable for computer. The vibrometer was located 376 mm from bearing. The laser was pointed horizontally, vertically and axial. Data were taken for various bearing conditions for speed of 400, 600, 800, 1000, and 1200 rpm. The extracted features covered mean, median, modus, root mean square, standard deviation, skewness, kurtosis, and beta kurtosis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003614_j.procir.2020.04.092-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003614_j.procir.2020.04.092-Figure3-1.png", + "caption": "Fig. 3. Motions of the DED nozzle and Deposition volume.", + "texts": [ + " Therefore, to respect the tool accessibility constraints of the DED process, it is necessary to modify the intersection part. In the proposed method, the intersection part and DED nozzle is modelled as \u03a6\ud835\udc56\ud835\udc56 and \u03a6\ud835\udc5b\ud835\udc5b by the level-set functions; Also, the deposited material is represented as \u03a6\ud835\udc53\ud835\udc53/\u03a6\ud835\udc56\ud835\udc56 , which can be calculated by min (\u03a6\ud835\udc53\ud835\udc53, \u2212\u03a6\ud835\udc56\ud835\udc56). As far as the collision detection is concerned, it is necessary to investigate the motion space (the union of motion) of the DED nozzle during new material deposition (see Fig. 3). In the proposed method, dilation is used to calculate the motion space of the DED nozzle. As one of the basic morphology image operations, dilation of a solid X by a structuring element \u2133 deposited as: \ud835\udc37\ud835\udc37(\u2133, \ud835\udc4b\ud835\udc4b) = \u22c3 \ud835\udc4b\ud835\udc4b\ud835\udc5a\ud835\udc5a \ud835\udc5a\ud835\udc5a\u2208\u2133 (9) where \ud835\udc4b\ud835\udc4b is a binary image in \ud835\udc38\ud835\udc38 (a Euclidean space), \u2133 is a structuring element (\u2133 \u2286 \u211d\ud835\udc51\ud835\udc51), \u222a is the union operation. In the study, the dilation function \ud835\udc37\ud835\udc37() is extended to represent the sweep of the DED nozzle by the deposition volume. Therefore, the input variables in Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002935_0954408920932358-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002935_0954408920932358-Figure1-1.png", + "caption": "Figure 1. Radiation structure model of inner pipe in spherical robot. (a) zero element; (b) one element; (c) two element.", + "texts": [ + " According to the design of electromagnetic power pipeline model, based on the principle of electromagnetic power generation, equation of induced electromotive force(EMF) under electromagnetic coupling will be derived, which is the theoretical basis of space layout of power generation pipeline structure. The concept of multi pipes structure model is put forward by referencing natural chemical molecular structure. Spatial topology models of robot The spherical robot model structures, such as the Tumbleweed Polar Rover of NASA2 and the internal and external drive of Xi\u2019an Electronic and Science University,20 are an integrated structure of single pipe. In order to facilitate the study, the ways of the pipe layout are named. As shown in Figure 1, they are called the zero element radiation structure, the one element radiation structure and the two element radiation structure. Considering the installation problem and stability, the topology node number of the pipe structure should be limited. In fact, the three element structure has not been found in the natural molecular structure. Because the structure will lead to the instability, its practicality and rationality are insufficient. Inspired by the chemical molecular structures which have good stability and space symmetry, this paper begins to study those structures including methane and pentane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000724_978-981-13-3305-7_190-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000724_978-981-13-3305-7_190-Figure2-1.png", + "caption": "Fig. 2. The tilt wing angle demonstration", + "texts": [ + " The tilt-wing mechanism of the target UAV is composed of stepper motors and corresponding structure made of carbon fibre material, whose main function is to make the wing rotate at a certain speed or stop at a certain deflect angle when UAV is in different stages of the transition mode. The dynamic model of the tilt-wing mechanism is: iw : \u00bc Xtilt \u00f02:5\u00de where iw represents the deflect angle of wing, whose value is 0 when the chord of wing is vertical to the axis of fuselage and becomes 90 when the wing\u2019s chord is parallel to the fuselage\u2019s axis, which is shown in Fig. 2, Xtilt denotes a rotational speed constant, whose value is 7:3 10 2 rad=s. The Propulsion Model. The propulsion model with respect to the combination of brushless rotors and fixed propellers can be estimated by the propeller model, which can be represented by the following equation: T \u00bc CtI 2 \u00f02:6\u00de M \u00bc CmI 2 \u00f02:7\u00de where I denotes the information about the speed of propellers, Ct and Cm denote the dimensionless thrust coefficient and torque coefficient, which can be obtained by the experiment of thrust and torque properties with respect to the rotor speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003185_012018-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003185_012018-Figure1-1.png", + "caption": "Figure 1, figure 2 and figure 3 illustrate the construction of the three four-spring electromagnetic harvesters. The following notations are used in the figures: 1 - coil; 2- rare earth magnets, 3- upper base, 4- steel plate, 5- steel springs, 6- lower base and 7 metal fastening. Springs with wire diameter ds = 0.63 mm, spring diameter Ds = 4.63 mm and length ls = 20 mm are used. The steel plates are the same size and with 2 or 4 fixed permanent NdFeB37 magnets, measuring 20x10x2 mm. The first harvester has two magnets and weighs m1=15 grams - Figure 1; the second is with 2 magnets, spaced from each other by 8 mm - Figure 2; and the third one weighs m2 = 30 grams and has two magnets in two places, spaced from each other by a distance of 8 mm - Figure 3. The coils under study have diameters Dc1 = 20mm and Dc1 = 40mm, thickness lc = 4mm and N1 = 600 and N2 = 1200 turns with conductor diameter dc = 0.1 mm.", + "texts": [], + "surrounding_texts": [ + "Sources, converting mechanical energy into electrical, such as the electromagnetic harvesters, are increasingly replacing the battery-type low power consuming power-supplying electronic devices [1, 2]. There are usually two types of magnet motion in relation to the coil in these harvesters. In the first case, the magnet moves in parallel to the coil [3] and in the second one - vertically with respect to it, often being located in its air gap [4]. Three different structures of four-spring electromagnetic harvesters are considered here, with two different masses, represented by steel plates with 2 or 4 magnets each and two different coils as well. The harvesters have a fixed coil attached to them, in parallel to which the rare earth magnets move. The aim of the present work is to study the effect of the construction parameters (such as the weight of the concentrated mass, the number of turns, the influence of the coil area, the volume of the magnets, and the size of the air gap) on the output electrical parameters of the studied harvesters. 2. Exposition TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 When applying a sinusoidally varying force with a resonant frequency, the mechanical system: \u201cmass (plate with permanent magnets) \u2013 springs\u201d begins to oscillate in parallel with the coil. Thus the magnetic flux passes differently through the coil and creates an alternating magnetic field in a different way, thus inducing alternating electromotive force. The studied harvesters are nonlinear mechanical oscillating systems. Their simulations made by ANSYS R19.1 take into account the fixture, the gravity effect of the plates with permanent magnets and the mechanical characteristics of the used springs. The horizontal deviation x of the mechanical system \u201cmass (plate with permanent magnets) \u2013 springs\u201d was obtained while modeling the four-spring electromagnetic harvesters, Figure 4. The magnetic field distribution of the three electromagnetic harvesters was obtained by means of FEMM 4.2. Figure 5 shows the magnetic field distribution for the third four-spring harvester with two spaced magnets in two places at zero horizontal deflection, and Figure 7 presents the distribution at maximum deviation. Figure 6 and figure 8 illustrate the magnetic flux density changes along the length of the harvester coil at zero and maximum horizontal deviation. From Figure 6 it can be seen that at zero horizontal deviation the normal magnetic flux density is zero, and at maximum deviation the maximum magnetic flux density Bmax is obtained. TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 The horizontal deviation at a sinusoidally changing force with a resonant angular frequency \u03c9 is also sinusoidal ( ) sinmx t X t\u03c9= (1) The horizontal oscillatory speed is equal to ( ) ( ) sin 2m d x t v t X t d t \u03c0\u03c9 \u03c9 = = + (2) The magnetic flux can be expressed by the change in the cross section of the coil \u0410(t), through which the maximum magnetic flux density Bmax passes ( ) ( ) max dA t \u0424 t B dt = (3) The change in the cross section \u0410(t) over time equals the horizontal deviation variation in the time x(t) along the diameter D\u0441 of the corresponding coil. ( ) ( ) \u0441 dA t D x t dt = (4) From (3) and (4), the magnetic flux through the coil is obtained ( ) ( )max \u0441 \u0424 t B D x t= (5) The induced electromotive force in the coil of the electromagnetic harvester is ( ) ( )d\u0424 t e t N dt = \u2212 (6) From (5) and (6) for the induced electromotive force for no-load mode, it is obtained ( ) ( ) max \u0441 d x t e t N B D dt = \u2212 (7) The amplitude of the induced electromotive force in the coil is a function of the amplitude of the oscillatory speed \u03c9Xm m max \u0441 m\u0415 N B D X\u03c9= (8) The angular frequency of the forced oscillations can be expressed by the frequency f 2 f\u03c9 \u03c0= (9) TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 From (8) and (9) for the amplitude of the induced electromotive force in the coil at a resonance it is obtained active load. In it, Rc and Lc denote the active resistance and the inductance of the electromagnetic harvester coil, \u0435c(t) - the induced electromotive force, and RL is the active load resistance. UL indicates the rectified voltage over the load resistance. The active power, in DC mode, is calculated using the amplitude of the induced no-load electromotive force in the coil at resonance, the voltage on the germanium diode UD and the parameters of the equivalent circuit (14). In the resulting expression, b denotes the attenuation coefficient, which is determined by the logarithmic attenuation decrement \u03b4 [5], where T is the oscillation periodic time and \u0415m(t) is the amplitude of the measured electromotive force 2b m\u03b4= , ( ) ( ) 1 ln m m E t T E t T \u03b4 = + (13)" + ] + }, + { + "image_filename": "designv11_80_0003513_j.promfg.2020.08.087-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003513_j.promfg.2020.08.087-Figure1-1.png", + "caption": "Fig. 1. Process procedure of HSB of extrusion", + "texts": [ + " Consequently, titanium alloys are widely used in the key force bearing components in airplane like frame, which have large section size and complicated shape. For example, the composite materials used in Boeing 787 aircraft account for more than 50% of the airframe weight, and the corresponding titanium alloy consumption is 15% [3]. To substitute forging process, hot stretch bending (HSB) is developed to form titanium airframe components in new commercial aircraft by the Cyril Bath Company [4]. The process procedure is presented in Fig.1. Firstly, the titanium profile is resistance heated to the predetermined temperature, and then it is stretched and wrapped against the die. After the extrusion has arrived at its final position, the position of clamps will maintain for a selected creep time to induce stress relaxation. Finally, the alloy is allowed to cool at a controlled rate. By using this process, the profile parts can be formed accurately with low manufacturing cost through improving material utilization and reducing NC machining quantity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001810_6.2020-2006-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001810_6.2020-2006-Figure5-1.png", + "caption": "Fig. 5: DIC and Numerical strain field in the 3-point bend tests", + "texts": [ + " Extraction of key material parameters from the stress-strain behavior, such as modulus and isotropic hardening parameters in both tension and compression, were obtained based on the results presented in Fig. 4. In turn the 3-point bend tests of the additively manufactured beams were simulated to validate the material model. As shown in Fig. 3, the force-displacement response for the numerical simulations were able to capture the experimental response well. Furthermore, analysis of the strain field as captured experimentally by DIC was well replicated in the numerical results as shown in Fig. 5 The next step in advancing a span-wise extending vehicle concept was to determine the expected loads to which the wing section would be subjected. Therefore it was necessary to select an appropriate airfoil. As mentioned -150 -100 -50 0 50 100 -0.15 -0.10 -0.05 0.00 0.05 0.10 0.15 S tr es s (M P a) Strain (m/m) 875-1 875-2 875-3 875-4 875-5 875-6 875-7 Fig. 4: Stress-strain behavior for the outer fibers of the RGD 875 beams subjected to 3-point bend testing D ow nl oa de d by U N IV E R SI T Y O F T E X A S A T A U ST IN o n Ja nu ar y 20 , 2 02 0 | h ttp :// ar c" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003753_s42835-020-00541-3-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003753_s42835-020-00541-3-Figure5-1.png", + "caption": "Fig. 5 Illustrating of zigzag leakage flux at condition 2", + "texts": [ + " Due to giant magnetic effect of ferromagnetic materials, the flux emitted by right end of PM2 will cross through tooth tip 2 to PM3 not linking the windings in slot. With rotating motor, zigzag leakage flux increases continuously. When q axis of PM overlaps with the center line of 1 3 tooth 2, zigzag leakage flux reaches its maximum. Thus, the zigzag leakage flux \u03a6\u03b41 can be calculated by where, \u03b1p is pole-arc coefficient, Q is number of slots, p is number of pole-pairs. 3.2 1 \u2264 < 1 + 2 The zigzag leakage flux of right end of PM2 that goes through tooth tip 2 to PM3 is decreasing. When the left end of PM3 overlaps with s3 as seen in Fig.\u00a05, the zigzag leakage flux is zero. Thus, the zigzag leakage flux \u03a6\u03b42 can be calculated by 3.3 1 + 2 \u2264 < 1 + 2 + 3 The flux emitted by PM2 will completely link the winding in slot through tooth 2 till the right end of PM2 overlaps with s3 as seen in Fig.\u00a06. Thus, the zigzag leakage flux \u03a6\u03b43 is nonexistent at condition 3, as follows. But, the air-gap leakage flux is existent, which can be obtained from (9). (10) 1 = p m p (11) 1 = Q \u2212 p 1 \u2212 p 2 (12) 2 = p m p (\u2212 + 2 1) (13) 2 = Q \u2212 p 1 \u2212 p 2 (14) 3 = 0 3 = p ( 1 \u2212 p ) 3" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003310_ess50319.2020.9160058-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003310_ess50319.2020.9160058-Figure1-1.png", + "caption": "Fig. 1 Proposed axial flux permanent magnet machine", + "texts": [ + " It is well known the vector control allows to inject a negative d-axis current for magnetic flux weakening. Nevertheless, injection of negative current in this technique may effect on the permanent magnets and their properties. Another way to control magnetic flux in machine is usage the hybrid excited synchronous machines that have two sources of excitation: permanent magnets and excitation winding. Permanent magnets create a fix magnetic flux. The axial flux permanent magnet machine assembly is depicted in fig. 1. In fig. 1 proposed axial flux permanent magnet machine topology has following components: 1 \u2013 stator core, made of roll electrical steel, 2 \u2013 copper one layer three phase winding, 3 \u2013 permanent magnets located on surface of rotor inductor and having the same polarity relatively to inductor, 4 \u2013 pole cores, which form and dimension are identical to permanent magnets, 5 \u2013 rotor inductor that provide the magnetic flux passing, 6 \u2013 ferromagnetic hub, 7 \u2013 excitation coil, 8 \u2013 shaft. Electric drive based on HESM allows to provide twozone motor speed control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-Figure4-1.png", + "caption": "Fig. 4. Selection method. The shape of the pushing plate is chamfered for convenience of explanation.", + "texts": [ + " (b) After extending node N1 to the maximum, the square shaft moves to the left in order to extend the node N2 so that the left edge of the square shaft S1 matches the left edge of the pushing plate P2. At this time, the node N1 is kept extended. The mechanism can extend the node N2 through the process from (c) to (e) by performing the same operation as above. By repeating this operation in the same way, the mechanism can extend multiple stages by using a linear motion mechanism which has a stroke of one node. Fig. 4 shows the details of the mechanism to select the pushing plate as well as the node and to move them linearly in the telescopic structure. The node Ni has di in diameter, and the pushing plate Pi has an arc of diameter di. There is a square at the center of the pushing plate that is the same size as the square shaft, and it can move along the square shaft axis. (i) to (iii) represent the flow to select and extend the pushing plate P1 as well as the node N1. (i) Control the position of the square shaft so that the left edge of the square shaft S1 matches the left edge of the pushing plate P1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002113_rdcape47089.2019.8979106-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002113_rdcape47089.2019.8979106-Figure2-1.png", + "caption": "Figure 2 Plus configuration", + "texts": [ + " The force on the Quadrotor due to gravity acts in the negative z direction. 197978-1-7281-2068-3/19/$31.00 \u00a92019 IEEE The architecture of a Quadrotor consists of rotors mounted on the end of arms in cross configuration frame. The most commonly used motor to drive these rotors are Brushless DC (BLDC) motors[7]. The thrust force generated by motor 1, 2, 3 and 4 is given denoted by F1, F2, F3 and F4 respectively as shown in figure 1. There are two main Quadrotor configurations: plus and cross-configuration. The diagram in figure 1 shows the cross-configuration[19]. Figure 2 shows the plus configuration whereas the figure 3 shows the cross configuration for a Quadrotor. The cross configuration is preferred and it is used in this paper. The mathematical modelling of a Quadrotor is done using the force moment balance and thus equations of motion are obtained for a Quadrotor can be written as[20]: (1) where the thrust is caused by motor i, is the torque produced by each motor is denoted by , moments of inertia is represented by \u2019s and m denotes the mass of the Quadrotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001658_icems.2019.8921632-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001658_icems.2019.8921632-Figure9-1.png", + "caption": "Fig. 9. Magnetic flux density visualization of ST-MG: (a) HMM, (b) FEA.", + "texts": [ + " The inner air-gap and outer air-gap magnetic flux density distributions of SM-MG and ST-MG solved by FEA and HMM are shown in Fig. 4 \u2013 Fig. 7. For the SM-MG, the magnetic flux density obtained by HMM and FEA almost coincide. Notably, a little large difference can be observed for the inner air-gap magnetic flux density of ST-MG, as shown in Fig.7. That is because an equivalent method is adopted to transfer a rectangle into two sectors. Besides, a magnetic field distribution for all parts in SM-MG and ST-MG calculated via HMM and FEA are shown in Fig. 8 and Fig.9. The magnetic field distribution trends are the same for HMM results and FEA results. The strips in HMM results are caused by the assumption that the permeabilities in radial direction is a constant, which further causes error in the magnetic field distribution prediction. In addition, it can be seen that the outer air-gap magnetic flux density distribution of SM-MG is almost the same with that of ST-MG. This means that the shapes of PMs on the inner rotor have little influence on the magnetic field distribution after the modulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001051_ccdc.2019.8832344-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001051_ccdc.2019.8832344-Figure3-1.png", + "caption": "Fig 3. The wing sections divided by the division method. Different sections shown in the figure above will generate different aerodynamic forces and moments under the influence of slipstream at different rotational speed, and the total aerodynamic forces and moments can be expressed as below:", + "texts": [ + " The 31th Chinese Control and Decision Conference (2019 CCDC) 1481 Authorized licensed use limited to: University of Exeter. Downloaded on May 05,2020 at 19:31:29 UTC from IEEE Xplore. Restrictions apply. The aerodynamic characteristics of slipstream is calculated by the one-dimensional momentum theory [10]. Besides, only the lifts and drags generated by the wing which is under the effect of slipstream are considered. Firstly, an area division method toward the whole wing should be used [11]. In our paper, the whole wing area in divided into nine sections, which are demonstrated in the Fig 3. ( )W W aero W ii B BF L DR= + (13) Wg B B aero Li i Di iM l L l D= + (14) where the lift and drag force of each section is shown as below: 21 2 =W i i i LivL S C\u03c1 (15) 21 2 =W i i i DivD S C\u03c1 (16) According to [12], the diameter of the slipstream varies with the axial distance behind the propeller discs, the relationship is given by: /p pd vd v= (17) where pv is the induced velocity of the propellers and v denotes the induced velocity behind the propeller discs, which is given by 2 2 / 1 (2 / ] ) ( ) [1 p p p x d x x v d v + = + (18) Then the wing and wing control surfaces area iS and aiS affected by the induced flow can be determined by the double integration method over the planform of the wing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure9-1.png", + "caption": "Fig. 9. Flux density of one step slotting model", + "texts": [], + "surrounding_texts": [ + "The proposed model with 6 poles and 48 slots reduces the cogging torque effectively. The significant improvement of performance for the proposed model, as shown in Fig. 11 can be identified by the smallest cogging torque in the beginning 15 tested points in the mechanical rotor degrees (00~150), and the last 15 tested point in the mechanical rotor degrees of (450~600). The maximum cogging torque of the proposed model is 0.007031968 Nm. While the maximum CT at the rotor rotates for the original model is 0.5830463 Nm, then we got the CT reduction as much 98.79%. Compare with the onestep slotting model. It is observed that the CT Reduction as much 93.90% with the maximum CT for this one step slotting is 0.556849257 Nm. It means that the proposed model is highly accepted in this research. At the beginning of rotation from motionless or at low speed, the original model and the model 1 need more mechanical energy to attain the same speed rotation compared with the proposed model (two-steps model). The fluctuation distribution of cogging torque for the proposed model. Authorized licensed use limited to: City, University of London. Downloaded on July 10,2020 at 12:04:30 UTC from IEEE Xplore. Restrictions apply. V. CONCLUSION AND REMARKS The influence of the slot opening width on the pole magnet area was investigated in the paper. From the simulation results, it can be concluded that the smaller crosssection area of magnet pole results in the decrement of cogging torque and air gap normal flux density in the InsetPMMs. Additional slotting in the magnet edge reduces the magnet pole cross-section and the cogging torque. The novelty of the proposed Model achieves to adjust the magnet pole arc and distance without changing the rotor diameter and stator construction." + ] + }, + { + "image_filename": "designv11_80_0003312_aim43001.2020.9158983-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003312_aim43001.2020.9158983-Figure2-1.png", + "caption": "Figure 2. The design of the plantar shoes.", + "texts": [ + " For each foot, the second metatarsal region, heel, and hallux are mainly bearing areas of the weight *Hao Liu is with the State Key Laboratory of Fluid Power and Mechatronic Systems, Zhejiang University, 310027 China (phone: +86- 571-87951271; fax: +86-571-87951941; e-mail: hliu2000@ zju.edu.cn). 978-1-7281-6794-7/20/$31.00 \u00a92020 IEEE 91 Authorized licensed use limited to: Cornell University Library. Downloaded on August 28,2020 at 14:59:15 UTC from IEEE Xplore. Restrictions apply. of the human body when people walk [11]. Flexi force sensors are installed in the plantar shoes of the exoskeleton, placed in the second metatarsal region (SMR) and the heel. The design of the plantar pressure shoes is shown in Fig.2. The pressure signals are converted into volt through a signal converting circuit. The plantar pressure signals and the joint angle signals are collected by the acquisition card of Ni company, with the model of usb6343 and the sampling frequency of 200Hz. Walking speed is determined by the step length and step rate. In healthy adults, there is a significant regression relationship between the step speed, length and rate [12]. For a fixed walking speed, there are multiple combinations of step length and step rate" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000996_ivs.2019.8813855-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000996_ivs.2019.8813855-Figure8-1.png", + "caption": "Fig. 8. Bodies, joints, spring and damper elements of the multibody model. Visceral mass is represented by a spring-mass-damper element on the torso.", + "texts": [ + " The rider model is equivalent to the model proposed by [23]. Simulated with the same parameters (proband 1) it predicts a much more damped pitch rate. By adjusting the arm damping ratio to one third of its original value, the measured resonance can be modeled with reasonable accuracy. Note that this pitch motion also constitutes coupling of vertical and longitudinal dynamics. Therefore also the bike and rider response to brake pulses is analyzed by measurement and simulation. For this purpose the full multibody model as shown in Fig. 8 is used. The hydraulic brake pressure, that is measured during the test ride, is converted to a brake torque and fed into the model. The results can be observed in Fig. 10. The rear wheel speed gives an adequate measure for the vehicle\u2019s forward velocity, because for this front brake maneuver, other than the front wheel speed signal it is not superimposed with slip. As expected the rear wheel speed reduces while brake pressure is applied. An interesting phenomenon occurs afterwards, when the speed increases again, even though no drive torque is applied", + " Elastic cantilever bending usually causes a continuous bending angle function \u03d5b(x) over the whole beam length. Dynamically this results in an infinite number of modes and resonance frequencies. For the bicycle fork there are also additional elastic elements like the headset, suspension and hub bearings. However, when observing the fork pitch rate \u03c9y,f in frequency range, it can be seen that there is a single dominating bending mode at 20 Hz. In the model this is approximated by a single revolute joint with torsional spring stiffness kb (see Fig. 8). In order to fit this parameter a measurement with a constant brake pressure is recorded. The best fit is achieved with kb = 120 Nm/deg. Additionally parameter variations from 50 % to 150 % of this value are displayed. Lower stiffness values cause smaller eigenfrequencies and more pronounced peaks, which is intuitive. Higher values in change do not cause higher frequencies. The reason for this lies in the brake pressure excitation, which has a broadband noise component that is bandlimited to 20 Hz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003739_s12239-020-0106-8-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003739_s12239-020-0106-8-Figure8-1.png", + "caption": "Figure 8. The model of modular deformable tire in ADAMS.", + "texts": [ + " Under braking condition, the supporting force T provided by the ground to the tire is transferred to the petalshaped connecting block, and the pressure generated by the piston fixed to the petal-shaped connecting block on the air cavity is the same as that under driving condition, so force analysis is not carried out here. P 195/75R14 92S tire is selected as the reference in this paper. The relevant data about this type of tire is used as the reference design value of modularized deformable tire and the model is modeled in SolidWorks (as shown in Figure 8), and then the model is imported into ADAMS. Next, dynamic simulation analysis of the force transmission characteristics of the tire in four driving, acceleration, deceleration and braking conditions will be carried out. The physical parameters of the tire are shown in Table 1. 4.1. Simulation Analysis of Force Transmission Characteristics of Modular Deformable Tires under Driving Conditions When the wheel is driving on the road surface, it is difficult to measure the pressure of the external wheel because of the friction between the external wheel and the ground" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000542_978-3-030-20131-9_261-Figure32-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000542_978-3-030-20131-9_261-Figure32-1.png", + "caption": "Fig. 32. Schematic diagram of the process of single-line straight-through weaving beaded cooling pad.", + "texts": [], + "surrounding_texts": [ + "As the alternate threading actions by the hand-weaving method mentioned above is very cumbersome, results in more complicated and low-efficiency to achieve the weaving works on the machine. In order to make the weaving of the beaded cool pad easy to implement on the machine, it is necessary to decompose and transform the weaving route extracted from the hand-weaving process to find a new weaving method. 3.1 Warp and Weft Automatic Weaving Method The warp and weft weaving method is widely used in the weaving of woven fabrics. In this method, the yarns of two systems, which are perpendicularly to each other, are interwoven to form a product according to a certain rule. Among them, the yarn of one system along the length direction is called a warp, and the yarn of one system along the width direction perpendicular to the warp is called a weft. When applied to the weaving of the beaded cooling pad, the yarn is regarded as a string. More specifically, the warp is called a warp string and the weft is named a weft string, as shown in Fig 6. Due to the beaded pad is woven by the \"warp and weft weaving\" method, the bead in the adjacent four pad units will fluctuate greatly, and the maximum fluctuation direction is perpendicular to the plane formed by the warp and weft strings. In order to reduce this fluctuation, a joint is added at the Method Research and Mechanism Design of Automatic Weaving\u2026 2541 intersection of the warp and weft strings. The joint is composed of two equal-diameter holes whose axes are perpendicular to each other and in the same plane. The outer shape is designed to be spherical for aesthetic appearance, so the joint is called a cross ball joint, as shown in Fig. 7. Since the cross-ball joint is added at the intersection of the warp and weft strings, the warp and weft strings won't be interlaced and the weaving process is simple. The string can be straightly passed through the bead and the cross ball joint. The weaving result is illustrated in Fig. 8. Fig. 6. Connection diagram of wrap and weft strings. Fig. 7. Schematic diagram of the cross-ball joint. Stereogram Section graph Fig. 8. Schematic diagram of warp and weft weaving beaded cool pad. 3.2 Lock Stitch Sewing Weaving Method In the sewing machine, there is a locked stitch consisted of two stitches intertwined together, and the interlacing point in the middle of the sewing material can be demonstrated in Fig. 9. Fig. 9. Diagram of the locked stitch. Fig. 10. Effect diagram of lock stitch sewing weaving method. In the lock stitch sewing weaving method, the vertical alignment beads are regarded as the sewing material. And the transverse arrangement beads are placed on the horizontal suture. The weaving result can be demonstrated in Fig. 10, in which the dotted line indicates the suture. In this method, it is firstly needed to insert the vertical alignment beads into a plurality of hooks. The number of hooks is one more than the number of columns of the beaded pad, the next weaving process can be illustrated in Fig.11 [12-13]. Welt string Warp string Stereogram Section graph Welt string Warp string S. Ouyang et al.2542 3.3 Single-line Straight-through Method The single-line straight-through method is a kind of weaving method that is easy to implement on the machine. Firstly, insert all beads that downlink line needed from one row of the beaded pad into the string, and the downlink line can be in any shape. Then the weaving process of every row of the beaded pad can be achieved by the threading movement of the uplink line. Finally, every row of the beaded pad can be joined together to form the desired b eaded pad. And the effect of this method is the same as that of hand-weaving method. For the beaded cooling pad LDm\u00d7p, each row has p units, so the number of beads required for the down line is 2p+1. The specific weaving steps are shown in Fig.12. And repeat steps in Fig.12 until the desired single row of beaded pads with units are woven, as shown in Fig. 13. Next each row of weaving beaded pad is repeated from the first step until the required number of beaded pad are completed, and then the single row of beaded pad are joined together to form the desired beaded cooling pad. To apply this weaving method on the machine, the machine can be designed in two parts, the first part is used to complete step 1 and the second part is used to complete steps 2 -7. It can be seen from the weaving process that the first part only has the downlink line threading, while the second part only has the uplink line threading. The whole process avoids the cumbersome action of repeatedly crossing the uplink and downlink lines. The beaded cooling pad woven by this method has a better stable and reliable string connection. As it ca n only be woven one row or one column at a time, it is easier to implement on a machine and has higher efficiency, providing an effective weaving method for designing and manufacturing a beaded cool pad automatic weaving machine. Method Research and Mechanism Design of Automatic Weaving\u2026 2543" + ] + }, + { + "image_filename": "designv11_80_0003990_ever48776.2020.9243015-Figure20-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003990_ever48776.2020.9243015-Figure20-1.png", + "caption": "Fig. 20. Rotor-bearing system, air duct system, and house of the prototype motor. (a) Rotor-bearing system. (b) Air duct system and house.", + "texts": [ + " Therefore, with the increase of stator lamination axial length, the influence of long end-winding axial length can be reduced in terms of torque density and efficiency. VI. EXPERIMENTAL VALIDATION A 180 krpm, 450W, 6-slot/2-pole HSPM motor with 2 coil-pitch windings has been optimally designed. The stator structure and winding configuration of the prototype motor are shown in Fig. 19. For high-speed operation, the rotor-bearing system, air duct system, and house of the prototype motor are designed and shown in Fig. 20. The FEM and measured phase back EMF waveforms of the prototype motor are compared in Fig. 21, and they have a good agreement. The current and voltage waveforms at 178 krpm are shown in Fig. 22. Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 22,2020 at 14:27:45 UTC from IEEE Xplore. Restrictions apply. Three 6-slot/2-pole HSPM motors with 1, 2, and 3 coil-pitch windings have been optimally designed by FEM. It shows that with a fixed Bmax, the coil-pitch has no influence on the optimal split ratio and thus three motors with different coil-pitches exhibit the same optimal split ratio" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure8-1.png", + "caption": "Figure 8. Stephenson1 mechanism with slider.", + "texts": [], + "surrounding_texts": [ + "The primary objective of this study was to identify the joint that producesthe highest error in a mechanism due to clearance. Simulations were carried out for all mechanisms described in the previous section by introducing clearances in different joints, one at a time. The effect of different speeds (of crank) and clearance sizes was also studied. Some important results are discussed here. The results for the four bar mechanisms have been summarized in figures 13, 14 and 15. These show that Joint 3 is the most critical joint and the values vary significantly for 0.02 mm but they are very similar for 0.5 mm. So, sensitivity to speed is high for low clearance. Figures 16, 17 and 18 show mean deviation at different joints of different inversions of the Watt chain at 0.1 mm radial clearance. The trend of all the curves is the same. Joint 3 is the most critical joint in all the three mechanisms followed by joints 2, 4 and 5, i.e., the criticality order here is: Joint 3[ Joint 2[Joint 4[ Joint 5. Joint 6 in figures 5 and 6 is the grounded joint which is least critical of them all. Figures 19 and 20 show the criticality graphs for Stephenson1 mechanism and Stephenson1 mechanism with slider, respectively. Both the graphs show the same trend where the deviation is maximum at the output link and decreases as one proceeds to the other end of the mechanism. A common trend which seems to emerge is that the first joint next to the crank, which has a link connected to ground, is the most critical joint. After that the criticality ranking goes towards the crank and then goes further down to the joints in order of their distances from the crank. We would examine the validity of this hypothesis in the rest of the study. Figures 21 and 22 show the criticality graphs for Stephenson2 mechanism (1) and Stephenson2 mechanism (2). The order of criticality for both the mechanisms is: Joint4 [ Joint3 [ Joint2 [ Joint5. This is as per the hypothesis stated earlier. The most critical joint is the joint next to the crank, which has a link connected to the ground (i.e., Joint 4). The next most critical joints are the ones towards the crank, i.e., Joint 3 and Joint 2. Then comes the joints which are left, i.e., Joint 5. Joint 6 is the grounded joint. The two graphs look very similar as the structures of mechanisms are the same. The deviations increase significantly with increase in speed, if the clearance is at Joint 4. The values in first graph are slightly higher than the corresponding values in the second graph. The overall conclusion is that deviations are more, if the output link is closer to the crank. Similar results are seen in eight and ten-bar mechanisms as shown in figures 23 and 24. The joint next to the crank is the most critical joint. It should also be noted that the grounded joints also show a trend. The grounded joint in the first loop has a significant effect but rest of the grounded joints have very little effect on the output. Error in output generally increases with increase in the crank speed (expect for an aberration shown in figure 19). This happens for all clearance sizes. However, the change in mean deviation is quite small even with a 5 times increase in speed from 100 to 500 rpm. Figures 25 and 26 show the variation of mean deviation with increase in speed when a clearance of 0.1 mm is at Joint 2 and Joint 3, respectively. It can be noticed that the deviations usually increase with speed. This means that a machine working at higher speeds is expected to produce more error than a machine working at lower speeds. Increase in radial clearance size brings an almost linear increase in the deviation values as seen in figures 27 and 28. It is also worth noting from figure 27 here that the deviation of slider is nearly 50% for 0.1 mm clearance (0.0501 is 50% of 0.1) and it goes on increasing to 80% for 0.5 mm clearance (0.3977 is 80% of 0.5). This happens because at lower values of clearance, the number of collisions between journal and bearing is more. Due to more number of collisions, the journal is sent back to the center again and again, and thus it remains at the center for a larger amount of time as compared to the case of large clearances. Thus, in larger clearances, the journal appears to remain closer to the walls of the bearing for an extended period and spends less time going to the diametrically opposite wall." + ] + }, + { + "image_filename": "designv11_80_0001394_isemantic.2019.8884342-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001394_isemantic.2019.8884342-Figure1-1.png", + "caption": "Fig. 1. Three Phase Switched Reluctance Motor", + "texts": [ + " In regenerative systems there are three methods of control, current control, angle position control and PWM control at voltage [4]. This paper will discuss regenerative systems in SRM, by using the PWM control method and setting the duty cycle, during the inductance decreasing using boost chopper circuit. The resulting voltage from boost chopper circuit will be greater than the battery voltage then the higher current can be delivered to the battery. II. RESEARCH METHODE An SRM has an iron laminates rotor with a salient pole while its stator has simple winding as depicted in Fig. 1. Torque in the SRM motor is generated by exciting the stator winding. When rotor and stator are on the aligned position, its inductance has maximum value, but the reluctance has minimum value. Meanwhile, under unaligned position, its inductance is low with high reluctance. To develop torque, excitation current must be given when the rotor in unaligned position. To control SRM, power electronics is required. Switched Reluctance motor can be controlled with several 978-1-7281-3832-9/19/$31.00 \u00a92019 IEEE converter topologies such as Asymmetric, C-Dump, R-Dump, and N+1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000055_012044-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000055_012044-Figure3-1.png", + "caption": "Figure 3. Geometry of computational domains with two kinds of shrouds.", + "texts": [ + " The windage power loss is also greatly affected by the volumetric flow rate around high speed gear. The power loss can be reduced by decreasing the volumetric flow rate. It is necessary to install the shroud for decreasing the volumetric flow rate around the gear. This study focuses on the influences of smooth and grooved shrouds on windage power loss for rotating speed ranging from 5,000rpm to 7,000rpm. The schematics of the high speed gears with smooth shroud and grooved shroud are shown in figure 3. The terms of c and h are respectively axial and radial distances between shroud and gear. The term c keeps constant at 4mm in this study. The grooved shroud is different than smooth shroud in that grooved shroud has grooves along the circumference. The cross-section of the groove is 0.5mm in width and 0.5mm in height. \u00a0 \u00a0 a. smooth shroud 5th AMMSE 2018 IOP Conf. Series: Materials Science and Engineering 473 (2019) 012044 IOP Publishing doi:10.1088/1757-899X/473/1/012044 Air flow around the gear teeth has been investigated numerically" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003762_2050-7038.12669-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003762_2050-7038.12669-Figure2-1.png", + "caption": "FIGURE 2 The equivalent model of the faulty 3\u03d5 IM in different frames; A, Stationary frame and B, rotational frame. IM, induction machine", + "texts": [ + "15 \u039brj j= ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi MsrdMsrq p Ieds\u2212\u03c4rP \u039brj j, \u00f014\u00de \u03a9sl = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi MsrdMsrq p \u03c4r \u039brj j Ieqs, \u00f015\u00de Te = pole 2 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi MsrdMsrq p Lr Ieqs \u039brj j, \u00f016\u00de \u03b8= \u00f0 \u03a9dt= \u00f0 \u03a9sl +\u03a9r\u00f0 \u00dedt, \u00f017\u00de where, the superscript \u201ce\u201d indicates the rotational frame. Based on (14) to (17), the vector control equations of the 3\u03d5 IM during OPF are obtained in the same way as the vector control equations of a healthy 3\u03d5 IM. The only difference is that Msrd is used instead of ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi MsrdMsrq p in the FOC equations of a healthy 3\u03d5 IM. After OPF condition in the 3\u03d5 IM, the machine model in the stationary frame is equivalent to an asymmetric 2\u03d5 IM with Nd and Nq turn numbers15 (see Figure 2A). In addition, using the presented unbalanced rotational matrix for the components of the current (Tis(\u03b8)), the machine model in the rotational frame is equivalent to a symmetric 2\u03d5 IM withffiffiffiffiffiffiffiffiffiffiffiffi NdNq p turn numbers15 (see Figure 2B). Based on Figure 2B, the stator d-q voltages in the rotational frame are written as (18: Ve ds Ve qs \" # = Ze ds 0 0 Ze qs \" # ffiffiffiffiffiffi Nd Nq s cos \u03b8\u00f0 \u00de ffiffiffiffiffiffi Nq Nd r sin \u03b8\u00f0 \u00de \u2212 ffiffiffiffiffiffi Nd Nq s sin \u03b8\u00f0 \u00de ffiffiffiffiffiffi Nq Nd r cos \u03b8\u00f0 \u00de 2 666664 3 777775 Isds Isqs \" # |fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl} Ieds Ieqs \" # : \u00f018\u00de Considering Ze ds Zs ds ffi NdNq N2 d = Nq Nd and Ze qs Zs qs ffi NdNq N2 q = Nd Nq , (18 is written as (19: Ve ds Ve qs \" # = Nq Nd Zs ds 0 0 Nd Nq Zs qs 2 664 3 775 ffiffiffiffiffiffi Nd Nq s cos \u03b8\u00f0 \u00de ffiffiffiffiffiffi Nq Nd r sin \u03b8\u00f0 \u00de \u2212 ffiffiffiffiffiffi Nd Nq s sin \u03b8\u00f0 \u00de ffiffiffiffiffiffi Nq Nd r cos \u03b8\u00f0 \u00de 2 666664 3 777775 Isds Isqs \" # : \u00f019\u00de As Vs ds =Zs dsI s ds, V s qs =Zs qsI s qs, and Zs ds Zs qs ffi N2 d N2 q , (19 can be simplified as (20: Ve ds Ve qs \" # = ffiffiffiffiffiffi Nq Nd r Zs dscos \u03b8\u00f0 \u00de Nq Nd Nd Nq 2 Zs qs ffiffiffiffiffiffi Nq Nd r sin \u03b8\u00f0 \u00de \u2212 Nd Nq Nq Nd 2 Zs qs ffiffiffiffiffiffi Nd Nq s sin \u03b8\u00f0 \u00de ffiffiffiffiffiffi Nd Nq s Zs qscos \u03b8\u00f0 \u00de 2 66664 3 77775 Isds Isqs \" # = ffiffiffiffiffiffi Nq Nd r cos \u03b8\u00f0 \u00de ffiffiffiffiffiffi Nd Nq s sin \u03b8\u00f0 \u00de \u2212 ffiffiffiffiffiffi Nq Nd r sin \u03b8\u00f0 \u00de ffiffiffiffiffiffi Nd Nq s cos \u03b8\u00f0 \u00de 2 666664 3 777775 Vs ds Vs qs \" # : \u00f020\u00de Thus, an unbalanced rotational matrix for the components of the stator voltage (Tvs(\u03b8)) can be achieved as (21): Tvs \u03b8\u00f0 \u00de= ffiffiffiffiffiffiffiffiffi Msrq Msrd r cos \u03b8\u00f0 \u00de ffiffiffiffiffiffiffiffiffi Msrd Msrq s sin \u03b8\u00f0 \u00de \u2212 ffiffiffiffiffiffiffiffiffi Msrq Msrd r sin \u03b8\u00f0 \u00de ffiffiffiffiffiffiffiffiffi Msrd Msrq s cos \u03b8\u00f0 \u00de 2 666664 3 777775: \u00f021\u00de Using (9), (12), (13), and (21), the stator d-q voltages of the 3\u03d5 IM during OPF in the rotational frame are written as (22: Tvs \u03b8\u00f0 \u00de Vs ds Vs qs \" # =Tvs \u03b8\u00f0 \u00de rs + LdsP 0 0 rs + LqsP \" # T\u22121 is \u03b8\u00f0 \u00deTis \u03b8\u00f0 \u00de Isds Isqs \" # +Tvs \u03b8\u00f0 \u00de MsrdP 0 0 MsrqP \" # T\u22121 \u03b8\u00f0 \u00deT \u03b8\u00f0 \u00de Isdr Isqr \" # " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001707_icems.2019.8921756-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001707_icems.2019.8921756-Figure8-1.png", + "caption": "Fig. 8. External damped sandwich structure structure", + "texts": [ + " The damped sandwich has an inner diameter of 180 mm and an outer diameter of 230 mm. The built-in damped sandwich structure is shown in Fig. 7. The external damped sandwich stator core is embedded in t the core, and the sandwich is embedded on both sides of the stator core. The thickness of the core is still 117mm. The parameters of the damped sandwich are: outer diameter is 230 mm, inner diameter d is 180 mm, thickness is 15 mm, and total thickness is 30 mm. The external damped sandwich structure is shown in Fig. 8 The built-in damped sandwiches are embedded in the middle of the stator core. The thickness of the damped sandwich is 10mm, 15mm, 20mm, 25mm, 30mm. Workbench software is used to calculate the first 100 natural frequencies of the five different damped sandwich thickness cores and the original stator core model, and the thickness of damped sandwich of stator1 to stator5 is 10mm to 30mm. The corresponding natural frequency graph is shown in Fig. 9. It is easy to find that the natural frequency of the stator core with the damped sandwich at the same order is lower than the natural frequency of the original core and increasing the thickness of the damped sandwich can effectively reduce the natural frequency" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002372_humanoids43949.2019.9035012-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002372_humanoids43949.2019.9035012-Figure3-1.png", + "caption": "Fig. 3. Figure of Agonist muscle and Antagonist muscle.", + "texts": [ + " In Section II, we will propose two control methods, antagonist modifier and agonist modifier. In Section III, we will show two experiments which evaluate these proposing methods. Finally, in Section IV, we will give a conclusion and future works of this research. Muscles are classified into two types, agonist muscles and antagonist muscles. In musculoskeletal humanoids, one joint is driven by multiple muscles like the human body. Those types are defined in relation to joint move direction, that is, for example, yellow arrow in Fig.3. Agonist muscles give force to drive joints in the desired direction. In Fig.3, blue muscle and green muscle are agonist muscles. It means 978-1-5386-7630-1/19/$31.00 \u00a92019 IEEE 601 that elbow joint is driven in the direction to the yellow arrow when these muscles pull wires. Antagonist muscles are stretched by joint movements. In Fig.3, red muscle is an antagonist muscle. It means that this muscle is stretched when elbow joint moves in the direction to yellow arrow. Antagonist modifier can be paraphrased as suppression of antagonist muscles. In an ideal robot model, antagonist muscles follow joint movement accurately. It means that tension of antagonist muscles are always zero and they have no looseness. However, in the real robot, differences with the ideal robot model cause too much tension of antagonist muscles. This tension opposes muscle movement and higher tension of agonist muscles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002173_icrera47325.2019.8996549-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002173_icrera47325.2019.8996549-Figure1-1.png", + "caption": "Fig. 1. Scanning electron micrograph of the thermal actuator. Two mirrored V-shaped beam stacks are connected to a lever beam. For the analytical model, the structure is separated into V-shaped beam stacks and a lever transmission with the spring stiffnesses kV, kL, respectively.", + "texts": [ + " Since the storage and transportation temperatures of refrigerator cargo or blood transfusion bags in clinical usage is critical, such a sensor is of advantage to ensure a proper cooling chain [22], [23]. The actuator itself is designed to be operated in a temperature range from \u221250 \u00b0C up to +100 \u00b0C. This range is determined and limited by the occurring mechanical stress due to the deformation of the structure. Finally, the paper is summarized in a conclusion in section VIII. The presented actuator design relies on V-shaped (\u201cchevron type\u201d) beams. Two packs of six 800 \u03bcm long beams in parallel with a tilting angle of \u03b3 = 4\u00b0 form one half of the actuator stack (Fig. 1). This beam stack has a mirrored counterpart which is shifted along the mirror axis (symmetry A). Both stacks are connected via coupling bars to a lever that moves in-plane depending on the change in temperature. The movable structure is fixed to a Si-substrate at four anchor regions at the corners of the structure. Mechanical stress energy is accumulated within the V-shaped beams due to differences in the thermal expansion of the active structure material (Ni), the substrate (Si), and temperature changes of the surrounding environment", + " A sort of hinge is formed by short flat springs in order to reduce the deformation energy of each V-shaped beam. These hinges have a width of 10 \u03bcm, whereas the rest of the V-shaped beam is 40 \u03bcm in width. The coupling bars are 475 \u03bcm in length and feature also a width of 10 \u03bcm. The offset doff between the beam stacks amounts 200 \u03bcm, and the lever exhibits a length of hL = 1105 \u03bcm with a width of 50 \u03bcm. The structure is 15 \u03bcm in height and consists of electroplated Ni on a Si-substrate. Table I lists all geometric parameters, where the individual variables can be found in detail in Fig. 1 and Figure 2. The actuator is designed for passive operation in a temperature range from \u221250 \u00b0C up to +100 \u00b0C, only relying on temperature changes of the surrounding environment. The geometry is separated into the V-shaped beam stacks and the lever transmission for the analytical model. The spring stiffnesses kV and kL of a single beam of the stack and the lever transmission, respectively, are calculated for a given deflection yd (Fig. 1). Finally, an equilibrium of the total energy due to the thermally induced expansion and the elastic energy of the deflected structure leads to the temperature dependent deflection at the tip of the lever xh (actuator stroke). A. Deflection The spring stiffnesses kV and kL as depicted in Fig. 1 are calculated in accordance to the static Euler-Bernoulli beam theory to be kV = E l1l2 I1 (l2 + l1) + 1 3 ( l3 2 I2 + l3 1 I1 ) (1) and kL = E ( 3 I3 l3 3 + A4 l4 ) I3 I4 l3 4 l3 3 + A4 I4 l2 4 3 + 3l2 4 l2 3 + A4 I3 l3l4 4 + I3 A3 3 l2 3 + 3A4 2A3 l3 l4 (2) where E is the Young\u2019s modulus, I1,2,3,4 are the second moments of area, A1,2,3,4 are the cross sectional areas, and l1,2,3,4 are the lengths of the individual beam sections (compare Fig. 2). The deflection of the tip of the lever xh in dependency of the deflection yd can be calculated with the length of the lever hL and the inclination w\u2032, resulting from solving the static Euler-Bernoulli beam theory, to be xh(yd) = [( l2 4 6I4 \u2212 l3 l4 A3 ) kL E + 1 l4 ] \u00b7 yd \u00b7 hL" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000753_1.5115895-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000753_1.5115895-Figure4-1.png", + "caption": "FIGURE 4.For 30 m/s free stream velocity, Velocity contour of NACA 0015 at (a) 8\u00b0 AOA;(b) 10\u00b0 AOA;(c)12\u00b0 AOA;(d)16\u00b0 AOA;(e)18\u00b0 AOA;(f)20\u00b0 AOA;", + "texts": [], + "surrounding_texts": [ + "The turbulent Spalart-Allmaras model has been chosen for the CFD simulation. The best location for the abrupt change in geometry is just prior to the region of separation. So, a backward facing step has been introduced just prior to the point of separation on the upper surface of the NACA 0015 airfoil and the lift and drag characteristics have been investigated by CFD analysis. The results are presented below. (a) (b) 050008-3 From 8\u00b0 to 10\u00b0 AOA, no separation can be observed. However, a separation can be observed near the trailing edge at 12\u00b0 AOA. From here on, the separated region grows and slowly moves towards the leading edge as the angle of attack increases. Near the trailing edge, a small separated region can be noticed at 14\u00b0 AOA. The separation point can be located where the x-wall shear becomes negative. In this case, from the CFD analysis it has been found to be at x/c = 0.83 from the leading edge. Stall conditions were reached at 18\u00b0 AOA. After which the flow over the upper surface of the airfoil becomes more separated. A very large increase in the separated region can be noticed when the angle of attack was increased from 18\u00b0 to 20\u00b0 due to exceeding the stall angle. 050008-4 From Figure 5 it can be observed that the point of separation slowly moves from the trailing edge to the leading edge at a uniform rate with the increase of angle of attack. In order to place the backward facing step, the angle of attack has been chosen to be 18\u00b0 and the free stream velocity to be 30 m/s. For 18\u00b0 AOA, the separation point has been found to be at x/c = 0.47 from the leading edge. The step has been placed at the point prior to the separation point; at x/c = 0.42 from the leading edge. 050008-5 (a) ( b) (c) FIGURE7.(a) Reduction in lift coefficient with increasing step depth.(b) Reduction in drag coefficient with increasing step depth. (c) Increase in lift to drag ratio with increasing step depth. From figure 7 (a) and (b), it is evident that both lift and drag reduction is linear until d/t = 0.07 and becomes irregular as d/t increases. However, the reduction in the lift is much smaller than the reduction in drag. As a result, from figure 7 (c) even though the lift is decreasing, for a larger decrease in drag, we get the higher lift to drag ratio with increasing step depth." + ] + }, + { + "image_filename": "designv11_80_0000647_s1028335819060041-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000647_s1028335819060041-Figure1-1.png", + "caption": "Fig. 1. Phase portrait of the unperturbed system.", + "texts": [ + " The amplitude of vertical oscillations of the platform is assumed small compared with the pendulum length (0 < , the oscillation frequency is arbitrary, and the parameter \u03b1 is considered close to its bifurcational value ( , where \u03b4 is a value around unity). The study is performed with classical and modern methods of nonlinear mechanics [6\u201312]. UNPERTURBED MOTIONS We represent the Hamiltonian function, Eq. (1), as (3) (4) The unperturbed system (when \u03b5 = 0) admits the integral of energy . The phase portrait of the unperturbed system is depicted in Fig. 1. The equilibrium \u03b5 \u2260 0 \u03c0 \u03a92 / \u03c0 \u03a92 / \u03b5 = \u03b1arccosq \u03c0 \u03a92 / \u03b5 1) \u03a9 \u03b1 = + \u03b5\u03b41 = + \u03b50 1,H H H = \u2212 + = \u2212 \u03b4 + \u03c4 2 0 2 1 1 1cos cos 2 , 2 4 ( cos )cos . H p q q H b q =0H h 8 position corresponds to the stable hanging pendulum; . In the equilibrium position p = 0, (or , which is equal from the physical viewpoint) corresponding to an unstable inversed pendulum, we have . For < 5/4 the pendulum oscillates periodically in the neighborhood of equilibrium with amplitude . For the pendulum is in the rotation mode" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001788_1350650119895192-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001788_1350650119895192-Figure3-1.png", + "caption": "Figure 3. Annular cavity with moved journal and bearing.", + "texts": [ + " In the cavitation model, the lubricant density is 800 kg/m3 for liquid and 0.55 kg/m3 for the vapour. The viscosity is 0.02 Pa s for the liquid and 0.0134mPa s for the vapour. The bubble number density for the Schnerr and Sauer Model is set to be 10.11 The saturation vapour pressure is 29185 Pa. As the geometric dimensions differ greatly in different directions of the oil film, a dynamic mesh method has been adopted for transit simulation in this paper to ensure fine mesh quality with the arbitrary relative motion between the journal and the bearing. Figure 3 shows a cross-section of a simplified plain journal bearing at any axial position. To describe the methodology of mesh movement clearly, the journal bearing structure shown has a larger clearance than in reality. O represents the coordinate origin. While Oj 0 and Ob 0 represent the current eccentricity positions after the journal and the bearing move from the upper position separately. For the point on the surface of the shaft and the bearing, i.e. A0(x1 0,y1 0) and B0(x2 0,y2 0), change their positions to A(x1,y1) and B(x2,y2)", + " When the journal and the bearing moves, the new position of P0 in the journal are defined by P marked in the figure. Then P0 P ! \u00f0 x; y\u00de can be defined as follows. P0P ! \u00bc P0B0 A0B0 A0A ! \u00fe A0P0 A0B0 B0B ! \u00f03\u00de In detail, the displacement of P is the linear interpolation of the relative motion between the shaft and the bearing. x \u00bc Ni N x1 x01 \u00fe 1 Ni N x2 x 0 2 y \u00bc Ni N y1 y 0 1 \u00fe 1 Ni N y2 y 0 1 ( \u00f04\u00de where Ni is the number of the radial reticulate layer in which P is located, and N is the total amount of radial reticulate layers in the annular clearance. Figure 3 also shows the grids for different bearing and shaft positions. From the corresponding side view of the journal bearing in the figure, no mesh distortion or numerical failure occur during the process. Besides, under the combination of large misalignment and eccentricity, the methodology can keep mesh angles optimal. Therefore, the proposed mesh movement method is suitable for the transient simulation, the hexahedron mesh of the oil film can change regularly according to the relative position between the shaft, and the bearing as the figure shows", + " For a flexible rotor-bearing system, the system motion equations can be given as follows. M \u20acx\u00fe C _x\u00fe Kx \u00bc FU \u00fe G \u00f05\u00de where x is the general coordinate of system, M, C and K are mass, damping and stiffness matrices, respectively, FU is the oil film force and G is the gravity of rotor supported by bearing. Figure 4 shows an axial section of the journal bearing. As the bearing clearance and the axial misaligned angle is too small relative to the bearing width, the movement of the shaft and the bearing could be described with the model as shown in Figure 3. The movement distance of the shaft is defined by five points as marked in Figure 4, which are obtained by solving the rotor dynamics equations in MATLAB. For an arbitrary point P between the Pk and the Pk\u00fe1, the movement of the point P can be given according to the following interpolation. xp \u00bc xk \u00fe lk lk\u00felk\u00fe1 xk\u00fe1 xk\u00f0 \u00de yp \u00bc yk \u00fe lk lk\u00felk\u00fe1 yk\u00fe1 yk\u00f0 \u00de k \u00bc 1, 2, 3, 4 ( \u00f06\u00de On the other side, the attitude angle of the bearing, marked as ( x, y, 0), can be solved by the following equations of motions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003989_ecce44975.2020.9235658-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003989_ecce44975.2020.9235658-Figure13-1.png", + "caption": "Fig. 13. 1/8 schematic of IPM motor B, unit [mm].", + "texts": [ + " The blue and red curves show without and with harmonic current injection, respectively. The red curve is mostly overlapped by the blue curve except 5000\u20138000Hz. It is seen that there is a significant peak at 6650Hz as high as 120dB without harmonic current injection. This frequency is corresponding to the mode 0 natural resonance of the stator core. It is found that the acoustic noise peak is significantly reduced by -27dB with the proposed small harmonic current injection. In this section, another harmonic current calculation example is shown for the improved IPM motor. Fig. 13 shows the cross section of the motor B, that is close to the IPM motor in the fourth generation Toyota Prius released in 2015. Compared with previous Prius motor, it has a special rotor structure and three rectangular permanent magnets are employed for one pole. Such special designs make the motor compact and efficient. 773 Authorized licensed use limited to: Tsinghua University. Downloaded on May 26,2021 at 09:21:34 UTC from IEEE Xplore. Restrictions apply. Let us define the origin of the rotor rotational position as shown in Fig. 13. As the rotor rotates, the radial force sum can be written as the function of the rotor position in the electrical angle, and the current which is composed of the d- and q-axis current id and iq . Let us assume the DC component of id and iq are id0 and iq0 respectively. Also, let us assume the AC component of id and iq are very small. By using the Taylor expansion, the radial force sum can be written as, Fsum(\u03b8, id, iq) =Fsum(\u03b8, id0, iq0) + \u2202Fsum(\u03b8, id0, iq0) \u2202id (id \u2212 id0) + \u2202Fsum(\u03b8, id0, iq0) \u2202iq (iq \u2212 iq0) (10) Let us consider only the 6th harmonic radial force in (10), then, Fsum(\u03b8, id, iq)6th = Fsum(\u03b8, id0, iq0)6th + \u2202Fsum(\u03b8, id0, iq0) \u2202id (id \u2212 id0)6th + \u2202Fsum(\u03b8, id0, iq0) \u2202iq (iq \u2212 iq0)6th (11) When the three-phase current is symmetrical sinusoidal waveform, id0 and iq0 are both DC component" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000811_romoco.2019.8787375-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000811_romoco.2019.8787375-Figure1-1.png", + "caption": "Fig. 1. Schematic of RTR manipulator", + "texts": [], + "surrounding_texts": [ + "I. INTRODUCTION\nMobile manipulator is a robotic system consisting of two subsystem: mobile platform and rigid manipulator. Such system has a possibility to manipulate different objects using a manipulation part and has an extended range of work, supported by the platform. However, one can observe dynamic interactions between subsystems, hence the proper control algorithm has to be applied in order to use full spectrum of mobile manipulator\u2019s capabilities.\nThree different tasks can be defined for robotics systems: point stabilization, trajectory tracking [9] and path following [2]. In the paper only path following task has been considered.\nFrom the robotics point of view, path is the curve parametrized by the curvilinear distance. Problem of path tracking has been elaborated many times, for instance for mobile robots [5], [1], [6] or mobile manipulators [7]. However, most of the papers present solution for following two-dimensional paths and it\u2019s practical usage is strongly limited.\nIn the paper solution for the mobile manipulator\u2019s path following problem has been proposed using orthogonal projection on the given 3D curve, based on Serret-Frenet frame moving along the path.\nThe paper is organized in the following way. Section II presents the general equation for modelling the RTR manipulator on unicycle platform. Description of robot\u2019s motion relative to path has been presented in Section III, while control problem has been defined in the Section IV. Section V contains the simulation results with some comment, while Section VI concludes the paper.\nII. MODELLING OF A MOBILE MANIPULATOR ON UNICYCLE PLATFORM\nLet\u2018s consider rigid RTR manipulator presented in Fig. (1) with fixed position of the first joint, mounted on the unicycle platform. State of such object can be described by a vector of platform\u2019s generalized variables qm and manipulator\u2019s variables qr, i.e. q = (qm,qr) T\n1 Chair of Cybernetics and Robotics, Electronics Faculty, Wroc\u0142aw University of Technology, ul. Janiszewskiego 11/17, 50-372 Wroc\u0142aw, Poland mateusz.cholewinski, alicja.mazur@pwr.edu.pl\nqm = ( x y \u03d5 \u03b11 \u03b12 )T \u2208 R5 (1)\nqr = ( \u03b82 \u03b83 )T\nwhere (x,y) are Cartesian coordinates of platform\u2018s mass centre defined in relation to a global frame X0Y0Z0, \u03d5 is unicycle\u2019s orientation, \u03b11 and \u03b12 are angles of rotation of left and right wheel, respectively, whereas \u03b8r is a vector of manipulator joint variables.\nPlatform is modelled as homogeneous rigid body with total mass Mp and moment of inertia Ip. Distance between manipulator\u2019s mounting point and center of platform\u2019s wheel axle is denoted as a, while symbol L is the half of the platform width. In turn, Ixx is the wheel\u2019s moment of inertia and Mk is the wheel\u2019s mass. Symbol R denotes radius of a wheel.\nThe first link of manipulator is modelled as the vertical joint of changeable length \u03b82. The second and third links are modelled as the rod of length l2, l3 and the mass m2, m3, respectively.\nA. Kinematics\nKinematics of mobile manipulator consist of two components: kinematics of holonomic manipulator and kinematics of nonholonomic platform [4].\n978-1-7281-2975-4/19/$31.00 \u00a92019 IEEE 203", + "1) Unicycle\u2019s kinematics: Constraints in skid-steering mobile platform\u2019s motion can be expressed in so-called Pfaffian form\nA(qm)q\u0307m = A(qm) x\u0307 y\u0307 \u03d5\u0307\n\u03b1\u03071 \u03b1\u03072\n= 0, (2)\nwhere A(qm), size of l \u00d7 5, is Pfaff matrix and l is the number of constraints. For a modelled mobile platform one can designate 2 constraints for lack of longitudinal slippage (one for each wheel) and one constraint for lack of lateral slippage.\n2) Manipulator: Kinematics of manipulator, denoted as \u03a6, describe the chain of transformations from the base coordinate system to the coordinate system associated with the end of manipulator. In the described situation it can be expressed in the following way\nA1 b = Rot(Z,0), (3) A2 1 = Trans(Z,\u03b82)Trans(X , l2)Rot(X , \u03c0\n2 ),\nA3 2 = Rot(Z,\u03b83)Trans(X , l3),\n\u03a6 = A1 bA2 1A3 2,\nwhere l2 and l3 are the length of second and third link respectively.\nManipulator kinematics should also include the transformation from the global coordinate system X0Y0Z0 to the coordinate system associated with the manipulator\u2018s base. It has following form:\nAb 0 = Trans(X ,x)Trans(Y,y)Rot(Z,\u03d5)Trans(X ,a), (4)\nwhere a is the distance between platform local coordinate frame XpYpZp and the place of manipulator installation.\nB. Dynamics\nFor mobile manipulator composed of the unicycle platform and holonomic manipulator general form of dynamics (not considering reaction forces associated with tire-ground contact) can be derived from d\u2019Alembert principle as follows\nM(q)q\u0308+C(q, q\u0307)q\u0307+D(q) = B(q)u+AT (qm)\u03bb ,\nor more in detail [4][ M11 M12 M21 M22 ]( q\u0308m q\u0308r ) + [ C11 C12 C21 C22 ]( q\u0307m q\u0307r ) + ( 0 D ) = [ B 0 0 I ]( um ur ) + ( AT \u03bb 0 ) . (5)\nSymbol M(q) denotes inertia matrix of mobile manipulator, matrix C(q, q\u0307) describes Coriolis and centrifugal forces, vector D(q) is a gravity forces vector. B(q) is the input matrix, indicating on which state variable the motor reacts. AT \u03bb are forces coming from nonholonomic constraints \u2013 from control point of view they are undesirable, making computations harder to be carried out.\nC. Model in auxiliary velocities\nSince the platform\u2019s velocity q\u0307m is in a null space of A(qm), it is always possible to find a vector of special auxiliary velocities \u03b7 \u2208 Rm, m = n\u2212 l (n is the number of platform state variables, n = 5 for unicycle platform, and l is a number of independent nonholonomic constraints). Thus, kinematics can be expressed in the form of driftless control system\nq\u0307m = G(qm)\u03b7 , (6)\nusing the G(qm) matrix, meeting condition A(qm) G(qm) = 0 where \u03b7 are the velocities having sense of angular velocities of the wheels. Set of auxiliary velocities for unicycle platform and generalized coordinates for holonomic arm will be called auxiliary coordinates of mobile manipulator.\nDynamics expressed in auxiliary coordinates take the following form\nQ\u2217 ( \u03b7\u0307\n\u03b8\u0308\n) +C\u2217 ( \u03b7\n\u03b8\u0307\n) + ( 0 D ) = B\u2217 ( um ur ) Q\u2217z\u0307+C\u2217z+ ( 0 D ) = B\u2217 ( um ur ) , (7)\nwhere [4]:\nQ\u2217 =\n[ GT M11G GT M12\nM21G M22\n] ,\nC\u2217 =\n[ GT (C11G+M11G\u0307) GTC12\nM21G\u0307+C21G C22\n] ,\nB\u2217 =\n[ GT B 0\n0 I\n] .\nz = (\u03b7 \u03b8)T (8)\n978-1-7281-2975-4/19/$31.00 \u00a92019 IEEE 204" + ] + }, + { + "image_filename": "designv11_80_0003742_icarm49381.2020.9195386-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003742_icarm49381.2020.9195386-Figure1-1.png", + "caption": "Fig. 1. Configuration of the orthosis.", + "texts": [ + " The next section is the derivation of the velocity transmission ratio between the device and the ankle joint. Then, the dynamics of the orthotic device is analyzed; accordingly, an angular tracking simulation is conducted. Finally, the hardware system of the orthosis is introduced. Based on the system, tracking of normal angle and angular velocity curve of the ankle joint are presented. 978-1-7281-6479-3/20/$31.00 \u00a92020 IEEE 582 Authorized licensed use limited to: University of New South Wales. Downloaded on November 15,2020 at 09:08:52 UTC from IEEE Xplore. Restrictions apply. Fig. 1 is the structure of the orthosis. It is mainly composed of the motor, gear box, bevel gear pair, main frame and bottom plate. Main frame, along with the motor and bevel gears, is fixed together with the shank by fixing belts; while the front foot is bound on the bottom plate. During walking, the assistant torque is provided by the power source. Specifically, the motor power is transmitted, in sequence, through bevel gears, Link 1, Link 2, Link 3 and finally to the front foot. The orthosis mainly has two functions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure2-1.png", + "caption": "Figure 2. Slider-crank mechanism.", + "texts": [], + "surrounding_texts": [ + "The most common computational tool used for design and analysis of multi-body mechanical systems is ADAMS [45, 46]. Various types of links and joints along with their properties like mass, location of center of mass, moment of inertia and degrees of freedom can be assigned for a mechanism in this tool [18]. The joints defined by this tool are ideal as they do not have any clearance or deformation. Therefore, to compare performance of the mechanisms with joint clearances, clearances at joints were made in ADAMS. The clearance at revolute joint between coupler and slider has been made by making a cylindrical hole in the slider cube and attaching a cylinder to the end of the coupler. So, we can change the size of the hole to set a clearance size. Similarly, the clearance at revolute joint between crank and coupler has been made by attaching a cylinder to the end of the crank and making a hole in the attached cylinder. 3.1 Input factors Various input parameters for modeling of mechanism in ADAMS are as follows: 3.1a Clearance size: For a standard journal-bearing of journal diameter 20 mm, the clearance size ranges from 0.02 mm to 0.08 mm. However, due to wear during operation and other environmental factors, the clearance can increase. So, in this research work, the clearance size has been taken in the range of 0.02 mm to 1 mm. 3.1b Crank speed: To cover a wide range of cases, the speeds range from 100 rpm to 3000 rpm. 3.1c Contact conditions: For the modeling of contacts, ADAMS uses the contact method based on the impact function: IMPACT-Function-Based Contact. In this method, the solver computes the contact force from the IMPACT function available in the ADAMS function library. The normal force of the contact has two components: rigidity and viscous damping. The component of rigidity is a function of the penetration d. The component of the viscous damping is a function of the speed of penetration. In this model the normal force of contact is given as: FN \u00bc Kdn \u00fe STEP d; 0; 0; dmax;Cmax\u00f0 \u00de _d; d[ 0 0; d 0 \u00f07\u00de 3.1d Value of K: The revolute joint is a case of contact between two cylinders (one inside the other). So, the contact should start with a line contact and then become a 2D rectangular contact. But the value of K in this case will not only depend on the material and geometrical property but also the stress distribution between the cylinders, which cannot be determined accurately unless it is a static case and the force is applied externally. So, the researchers solved this problem by stating that the line contact in the revolute joint will only be present for two cylinders aligned with extreme precision. Also, a uniform force distribution over the length of the joint is not possible in real life conditions. Moreover, the forcedeformation diagrams for both spherical and cylindrical impact force models were studied in the literature [1\u20134] and it was found that the spherical and cylindrical forcedeformation diagrams are reasonably close. Based on these studies, we used the Hertzian contact force law between two spheres with the different parameters defined in Eq. (8). K \u00bc 4 3p hi \u00fe hj R1=2;R \u00bc RiRj Ri \u00fe Rj ; hk \u00bc 1 v2k pEk ; k \u00bc i; j \u00f08\u00de Ri, mi and Ei represent respectively the radii of the cylinders, the Poisson\u2019s ratio and the modulus of elasticity for element i. For clearance = 0.02 mm, journal radius = 10 mm and bearing radius = 10.02 mm E = 2.07*105 N/mm2; m = 0.29 Putting these values in the equation we get K = 3.37*105 N/m1.5 Similarly, for clearance = 0.1 mm, K = 3.377*105 N/m1.5 For clearance = 0.5 mm, K = 3.4*105 N/m1.5 3.1e Value of n:The value of n is usually taken to be 1.5 for metallic contacts. So, n = 1.5. 3.1f Value of damping coefficient: In ADAMS, the instantaneous damping coefficient is a cubic step function of the penetration given as: STEP d; 0; 0; dmax;Cmax\u00f0 \u00de \u00bc 0; d 0 Cmax d dmax 2 3 2 d dmax ; 0\\d\\dmax Cmax d dmax 8 >>< >>: \u00f09\u00de The value of Cmax should be approximately 1 percent of the value of K. dmax = 0.01 mm 3.2 Output factors Two factors were used as parameters for comparing the kinematic performance of different mechanisms, either displacement of the slider or angular rotation of the rocker attached to ground." + ] + }, + { + "image_filename": "designv11_80_0003112_j.mechmachtheory.2020.103903-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003112_j.mechmachtheory.2020.103903-Figure6-1.png", + "caption": "Fig. 6. Fully parallel 5-DOF 4-coupled-Cartesian-manipulator with active moving-base (base-link L Bs ).", + "texts": [ + " 10 ( R P PP U ) ( P P P U ) \u2217 Extension to family of coupled-Cartesian-manipulators to investigate in further research \u2217 Enables rotation control of angle \u03b8T Z around common-link L T longitudinal \u02c6 Z T axis and D) In Sm illustrate manipulators with intersecting-revolute-axes. Each figure illustrates a stand-alone manipulator, with its own unique features that are summarized in Table 2 . The progression of Figs. 1\u20136 illustrates the hierarchical composition of the 5-DOF 3 T 2 R fully parallel mechanism of Fig. 6 . Two serial Cartesian manipulators Fig. 1 form the two hybrid serial- parallel manipulators of Figs. 3 , 4 which together form the single fully parallel 4-DOF 2 T 2 R manipulator Fig. 5 . Adding an active moving-base (base-link) extends it to the 5-DOF 3 T 2 R fully parallel manipulator Fig. 6 . To make it easier to see the mechanisms, only one out of four possible vertical supports L w is shown in Figs. 1\u20136 A. In practice, these mechanisms would typically have four vertical supports and redundant prismatic P 2 joints to evenly balance loads as shown in Figs. 6 B, 8 . The mechanism in Fig. 1 D represents a conventional ( P P P RRR ) serial Cartesian manipulator with a 3-DOF spherical wrist. Stuart [53] reports the manipulators in Figs. 2 , 8 . The manipulators in Figs. 3\u20136 , 9 , 10 are novel", + " This allows coaxial bearings R 1 , R 2 to be spaced far apart from each other, along the common-link L T , as seen in Fig. 5 A, to support high moment loads with high moment stiffness. The load support link in solid model In Sm Fig. 5 D supports external loads applied to the common-link L T and the weight of the links and joints along the common-link L T . A redundant passive ( PPRP ) limb connects the load support link to the workspace link L W . Its revolute R joint allows the load support link to swivel with changes in orientation of the common-link L T . Fig. 6 fully parallel 5-DOF 4-coupled-Cartesian-manipulator with active base-link L Bs . The two parallel-connected manipula- tors In Sc , In Sm in Fig. 6 A, 6B control the 5-DOF position and orientation of common-link L T relative to base-link L Bs . One of the prismatic joints along L W is fixed and the other three prismatic joints along L W are passive to accommodate changes in distance, in the \u02c6 Z W W direction between the limbs, as the common-link L T orientation changes. Joint topology for the fully par- allel 4-coupled-Cartesian-manipulator in Fig. 6 A is (P P UR )3( PP P U)( P ) for 5-DOF control with an active base-link L Bs actuated by an active prismatic ( P ) joint. Notice that only one prismatic P joint is active per limb indicating that the manipulators in Fig. 6 are fully parallel-connected. The five active prismatic P joints, with linear positions x A 1 A 1 , x C1 C1 , x A 2 A 2 , x C2 C2 , z W A 2 are identified with short arrows in schematic In Sc Fig. 6 A. 3. Mechanical design synthesis of a 4-coupled-Cartesian-manipulator Drawing In Sm Fig. 6 B is a solid model of a 5-DOF 3 T 2 R fully parallel 4-coupled-Cartesian-manipulator based on commercially available mechanical components, with redundant prismatic P 2 joints and joint topology (P 2 P UR )3(P 2 P 2 P U)( P 2 ) . It has a spindle attached to the common-link L T (moving-platform) for 5-axis machining applications with large workspace volume. The manipulator in Fig. 6 B is limited to a polar tilt angle | \u03b8p | \u2264 47 \u25e6 of the common-link L T longitudinal \u02c6 Z T axis. However, since machining tools such as end mills cut both on their sides and ends, the manipulator can contour surfaces over a range of polar tilt angles | \u03b8p | \u2264 137 \u25e6. The structural and mechanical components and overall layout of the manipulator are designed for stiffness to resist deflection from interactive machining forces applied to the spindle. The animation video file, Xactuator.mp4 in supplementary materials, shows how the manipulator in Fig. 6 B moves. Please cite this article as: P. Wiktor, Coupled Cartesian manipulators, Mechanism and Machine Theory, https://doi.org/10. 1016/j.mechmachtheory.2020.103903 12 P. Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx application of the mechanism based on coupling two sets of Cartesian manipulators together, with each set having parallelrevolute-axes. The Cartesian manipulators are coupled to the common longitudinal \u02c6 Z T axis of the common-link L T through nested revolute orthogonal RR joints with intersecting axes", + " The two sets of Cartesian manipulators, with parallel-revolute-axes, are orthogonal to each other and are coupled together, along the common-link L T longitudinal \u02c6 Z T axis, by a revolute R joint to accommodate the parasitic-twistangle \u03b8BD n Z . The limbs of the two sets are interleaved with each other, along the common-link L T longitudinal \u02c6 Z T axis, so that rotary bearings for the common revolute R 1 , R 2 joint can be spaced far apart from each other to support high moment loads. Solid model Fig. 6 B illustrates the importance of synthesizing parallel mechanism topologies that are conducive to construction based on fundamental mechanical design principles, Table 3 . Please cite this article as: P. Wiktor, Coupled Cartesian manipulators, Mechanism and Machine Theory, https://doi.org/10. 1016/j.mechmachtheory.2020.103903 P. Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx 13 Design simplicity is a guiding principle for the coupled-Cartesian-manipulators. Occam\u2019s razor, \u201cEntities should not be multiplied without necessity\u201d, may be interpreted here as, \u2018parallel mechanisms should not have more components than nec- essary\u2019", + " (32) cannot be rearranged, by alternative reordering of rows and columns [26] , into diagonal, triangular or block-diagonal matrices, the 4-coupled-Cartesian-manupator is not decoupled from the perspective of its geometrical Jacobian matrix. Singularities. The Jacobian J matrix Eq. (32) loses rank at | \u03b8T x | = 90 \u25e6 or | \u03b8T y | = 90 \u25e6. At these angles the manipulator loses its angular velocity y W \u03c9 T or x W \u03c9 T coordinate degrees-of-freedom. However the mechanical components of actual implementations of coupled-Cartesian-manipulators typically prevent these extreme angles from being reached in practice. For example, the workspace of the manipulator in Fig. 6 B is limited to polar tilt angle | \u03b8p | \u2264 47 \u25e6 due to the geometry of the nested intersecting revolute RR joints. So the manipulator is free of singularities in its workspace. 4.6. Parasitic-twist-angle \u03b8BD n Z between links L Bn , L Dn In general, links L Bn , L Dn , L T rotate relative to each other by parasitic-twist-angle \u03b8BD n Z around their common \u02c6 Z W Bn , \u02c6 Z W Dn , \u02c6 Z W T axis as the common-link L T orientation changes. The rotation matrix R Dn Bn expresses the orientation of link L Bn relative to link L Dn " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000025_s1068366618060144-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000025_s1068366618060144-Figure2-1.png", + "caption": "Fig. 2. Dependence of the excess pressure in the slit-like clearance on the polar angle \u03d5 for an isothermal f low at different eccentricity ratio values calculated by Eq. (7) (\u03a91 = 100 s\u20131; k = 0): (1) \u03b4 = 0.1, (2) \u03b4 = 0.3, and (3) \u03b4 = 0.5.", + "texts": [ + " (2), the pressure distribution depends in this case according to relation (6) not only on \u03a91, \u03d5, \u03b4, and \u03b5 but also on the distribution of the temperature T in the bearing clearance. ( ), \u03b4kS \u03d5 ( ) ( ) ( ) 0 , 1 cos 1,2,3 .k k kdS \u03d5 =\u03d5\u03d5 \u03b4 = \u2212 \u03b4 \u03d5\u222b \u03d5 \u03d5\u2202 \u2202\u2202 \u2202\u2202\u03c1 \u2202\u03c1\u03bc = \u03bc = + = \u2202 \u2202\u03d5 \u2202 \u2202\u03d5\u2202 \u2202 22 2 2 2 2 1 1, , 0,r rV VV Vc c r r r rr r ( ) ( ) ( ) 2 2 1 1 0 2 32 2 0 6 2(1 ), , , 2 r S S RT \u239b \u239e\u03bc\u03a9 \u2212 \u03b4\u03c1 \u03d5 \u2212 \u03c1 = \u03d5 \u03b4 \u2212 \u03d5 \u03b4\u239c \u239f\u03b5 \u03b3 + \u03b4\u239d \u23a0 ,p RT= \u03c1 ( ) ( ) ( ) 0 2 2 1 1 2 32 2 0 6 2(1 ), , , 2 p RT r T S S T \u03d5 = \u03c1 \u239b \u239e\u03bc\u03a9 \u2212 \u03b4+ \u03d5 \u03b4 \u2212 \u03d5 \u03b4\u239c \u239f\u03b5 \u03b3 + \u03b4\u239d \u23a0 L OF FRICTION AND WEAR Vol. 39 No. 6 2018 In an isothermal case at T = T0, Eq. (6) takes the form (7) For comparison, Fig. 2 shows the curves of the excess pressure distribution in the sliding bearing clearance constructed by Eq. (7) depending on the polar angle \u03d5 at different values of the eccentricity ratio \u03b4. From the graphs in Figs. 1 and 2, one can see that, in an isothermal case, the pressure distribution calculated by Eq. (7) differs from solution (2) only by the maximum value amplitude. As for the rest, they are identical to each other. The qualitative differences between two approaches to the description of the pressure distribution become apparent only considering changes in the temperature due to dissipative heating of the moving medium" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002002_robio49542.2019.8961771-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002002_robio49542.2019.8961771-Figure5-1.png", + "caption": "Figure 5. Simulation analysis of the continuum robot affected by external load. (a) The 3D model of bending continuum robot with constant curvature, The starting point of the robot is fixed end, the end is subjected to external loads (b) Deformation with 0.1N external load, (c) Deformation with 0.2N external load, (d) Deformation with 0.3N external load.", + "texts": [ + "1 connection disk, each two adjacent connecting disks are defined as one set of module according to the method shown in the Figure 4. There are 11 groups in total. The purpose of dividing modules is that the NiTi columns of each connecting disk are interlaced. If the continuum robot bends in a certain plane, only the columns between each module will bend. The concrete structure and bending motion mode are further elaborated in [21]. The 3D model of bending continuum robot with constant curvature is built as shown in Figure 5(a). The model is imported into the simulation software for force analysis. The starting point of the robot is fixed end, the end is subjected to external loads of 0.1N, 0.2N and 0.3N, respectively. The shape change of the continuous robot is shown in Figure 5(b)(c)(d). From the figure, we can draw a preliminary conclusion that the curvature from the end to the middle of the continuum robot has changed obviously. In order to get more detailed curvature change values, as shown in Figure 6, we use image recognition to analyze the curvature of the continuum robot under four different conditions(no external load, 0.1N, 0.2N and 0.3N external load), and mark the center of mass of each module with red dots . In the way of [1], a concise method for image recognition to obtain the shape of the continuum robot is proposed, which is a shape reconstruction algorithm for a section of arc determined by three points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003185_012018-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003185_012018-Figure4-1.png", + "caption": "Figure 4. Deviation of the mechanical system.", + "texts": [ + " Thus the magnetic flux passes differently through the coil and creates an alternating magnetic field in a different way, thus inducing alternating electromotive force. The studied harvesters are nonlinear mechanical oscillating systems. Their simulations made by ANSYS R19.1 take into account the fixture, the gravity effect of the plates with permanent magnets and the mechanical characteristics of the used springs. The horizontal deviation x of the mechanical system \u201cmass (plate with permanent magnets) \u2013 springs\u201d was obtained while modeling the four-spring electromagnetic harvesters, Figure 4. The magnetic field distribution of the three electromagnetic harvesters was obtained by means of FEMM 4.2. Figure 5 shows the magnetic field distribution for the third four-spring harvester with two spaced magnets in two places at zero horizontal deflection, and Figure 7 presents the distribution at maximum deviation. Figure 6 and figure 8 illustrate the magnetic flux density changes along the length of the harvester coil at zero and maximum horizontal deviation. From Figure 6 it can be seen that at zero horizontal deviation the normal magnetic flux density is zero, and at maximum deviation the maximum magnetic flux density Bmax is obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002507_ropec48299.2019.9057082-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002507_ropec48299.2019.9057082-Figure1-1.png", + "caption": "Fig. 1. A quadrotor with slung load.", + "texts": [ + " The quadrotor dynamics are under\u2013actuated and nonlinear due to the gyroscopic effects of the rotations of the rigid body and of the propellers. Besides, the following assumptions are considered: - The mass center is considerate the center geometric of the quadrotor. - The origin point in the earth coordinates is coincident to the origin point in the quadrotor coordinates - The cable that connect the slung load to the quadrotor is considerate rigid. Let us indicate with RC(O, e1, e2, e3), R\u0393(\u2126, \u03b51, \u03b52, \u03b53) the frames fixed with the Earth and the quadrotor, respectively, with \u2126 coincident with the center of mass of the quadrotor, see Fig. 1. The quadrotor absolute position in RC is described by p = (x, y, z)T , whereas its attitude is described by the Euler angles \u03b1 = (\u03c6, \u03b8, \u03c8)T , where \u03c6, \u03b8, \u03c8 \u2208 (\u2212\u03c0/2, \u03c0/2) are the roll, pitch and yaw angles, respectively. The sequence 1\u20132\u20133 has been considered [12]. Moreover, v = (v1, v2, v3)T , \u03c9 = (\u03c91, \u03c92, \u03c93)T are the linear and angular velocities of the center of mass of the quadrotor, expressed in R\u0393. The Euler-Lagrange equation is: d dt [ \u2202L \u2202q\u0307 ] \u2212 \u2202L \u2202q = ( Fext \u03c4ext ) (1) The Lagrangian system is represented as: L = T \u2212 U , where T represents the kinetic energy of the system and U the potential energy. The quadrotor position is considerate as xD, yD, and zD, while the slung load position are xL, yL, zL and are determined by a trigonometric analysis, as shown in the Figure 1. In this article with si = sin(i) and cj = cos(j), i, j = \u03c6, \u03b8, \u03c8, \u03b1, \u03b2. The slung load position is: xL = lLs\u03b2 c\u03b1 yL = lLs\u03b2 s\u03b1 zL = lLc\u03b2 . (2) The dynamic of slung load system is: x\u0307L = lL(\u03b2\u0307c\u03b2 c\u03b1 \u2212 \u03b1\u0307s\u03b2 s\u03b1) y\u0307L = lL(\u03b2\u0307c\u03b2 s\u03b1 + \u03b1\u0307s\u03b2 c\u03b1) z\u0307L = \u2212lL\u03b2\u0307s\u03b2 . (3) Then, the Lagrangian system is: L = 1 2 mD(x\u03072 D + y\u03072 D + z\u03072 D) + 1 2 mL ( (x\u0307D + x\u0307L)2 + (y\u0307D + y\u0307L)2 + (z\u0307D + z\u0307L)2 ) + 1 2 \u03c9TJD\u03c9 + 1 2 JL\u03b3\u0307 2 \u2212mDgzD \u2212mLg(zD \u2212 zL) (4) where, JD is the quadrotor inertial matrix, JL is the slung load inertial and \u03c9 is the angular speed rotation matrix as: \u03c9 = \u03c6\u0307 0 \u2212\u03c8\u0307s\u03b8 0 \u03b8\u0307c\u03c6 \u03c8\u0307c\u03b8 s\u03c6 0 \u2212\u03b8\u0307s\u03c6 \u03c8\u0307c\u03b8c\u03c6 The variables considered in the quadrotor-slung load system are: q = [xD yD zD \u03c6 \u03b8 \u03c8 \u03b1 \u03b2 \u03b3 ] and the derivative are: q\u0307 = [x\u0307D y\u0307D z\u0307D \u03c6\u0307 \u03b8\u0307 \u03c8\u0307 \u03b1\u0307 \u03b2\u0307 \u03b3\u0307 ] Solving the equation (4) became" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003614_j.procir.2020.04.092-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003614_j.procir.2020.04.092-Figure10-1.png", + "caption": "Fig. 10. Modified intersection part by considering the nozzle collision problem of DED process", + "texts": [ + " 9 shown, the used part is transformed from \u03a6\ud835\udc62\ud835\udc62 to \u03a6\u0303\ud835\udc62\ud835\udc62 by the optimal transformation matrix obtained by the proposed algorithm and then the intersection part \u03a6\ud835\udc56\ud835\udc56 is obtained. The result of the optimal transformation is shown in Table 3. After obtaining the intersection part, the next step is modifying the intersection part from \u03a6\ud835\udc56\ud835\udc56 to \u03a6\u0303\ud835\udc56\ud835\udc56 by respecting different AM processes. In the DED process, the material deposition nozzles have collisions with the part. The intersection part is modified from \u03a6\ud835\udc56\ud835\udc56 to \u03a6\u0303\ud835\udc56\ud835\udc56 by the implementation of Eq. (9)-(11). The result is shown in Fig. 10. In terms of the PBF process, the intersection part is modified automatically with respect powder recoater collision problem. The result is shown in Fig. 12, and the level-set function of the optimized intersection part is \u03a6\u0303\ud835\udc56\ud835\udc56 = min{ \u03a6\ud835\udc56\ud835\udc56, 180 \u2212 \ud835\udc65\ud835\udc65, \ud835\udc65\ud835\udc65 ,120 \u2212 \ud835\udc66\ud835\udc66 , \ud835\udc66\ud835\udc66 ,20 \u2212 \ud835\udc67\ud835\udc67 , \ud835\udc67\ud835\udc67}. The last step is individual features extraction. As shown in Fig. 12, a level set represented subtractive manufacturing volume is \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34. \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 is converted to \ud835\udc35\ud835\udc35\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 by the CSG to Brep conversion algorithm. Then, individual SFs (SF1 and SF2) are extracted by the machining feature recognition algorithm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003701_s00170-020-05818-5-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003701_s00170-020-05818-5-Figure2-1.png", + "caption": "Fig. 2 The schematic of the multiple reflection inside keyhole", + "texts": [ + " The software FLOW-3D was employed in the analysis, and the physical properties of 5182 Al alloy used in different plate thickness are shown in Table 1. The free surface is determined by researching the function of the volume ratio of the fluid in the grid cells. As long as the value of fraction of fluid in the total volume in the grid (F) is between 0 and 1, the grid cell is considered as a free surface area. The volume of fluid (VOF) method in FLOW-3D software can be determined as [22] \u2202F \u2202t \u00fe \u2207 V * F ! \u00bc 0 \u00f01\u00de The laser heat source can be calculated by the following equation [23]: Q \u00bc P \u03c0rb2 exp \u2212 r2 rb2 \u00f02\u00de Figure 2 is the schematic of the multiple reflection inside keyhole, and the Fresnel absorption ratio of laser beam can be expressed as follows [24, 25]: \u03b1 \u03c6 N\u00f0 \u00de \u00bc 1\u2212 1 2 1\u00fe 1\u2212\u03b5cos\u03c6 N\u00f0 \u00de 2 1\u00fe 1\u00fe \u03b5cos\u03c6 N\u00f0 \u00de 2 \u00fe \u03b52\u22122\u03b5cos\u03c6 N\u00f0 \u00de \u00fe 2 cos\u03c6 N\u00f0 \u00de 2 \u03b52 \u00fe 2\u03b5cos\u03c6 N\u00f0 \u00de \u00fe 2 cos\u03c6 N\u00f0 \u00de 2 0 B@ 1 CA m \u00bc 1 2 1\u00fe 1\u2212\u03b5cos\u03c6\u00f0 \u00de2 1\u00fe 1\u00fe \u03b5cos\u03c6\u00f0 \u00de2 \u00fe \u03b52\u22122\u03b5cos\u03c6\u00fe 2 cos\u03c6\u00f0 \u00de2 \u03b52 \u00fe 2\u03b5cos\u03c6\u00fe 2 cos\u03c6\u00f0 \u00de2 ! \u00f03\u00de The adiabatic bubble model has been loaded expound the keyhole-induced porosity formation. It is generally considered that the pressure and temperature in the entire bubble are homogeneous and compatible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003319_s10514-020-09938-5-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003319_s10514-020-09938-5-Figure2-1.png", + "caption": "Fig. 2 Schematic of the triple pendulum representing a leg of the orthosis-dummy system", + "texts": [ + " For the purpose of identification of the complete orthosis-dummy system, the triple pendulummodel is identified separately from the ankle stiffnessmodel since linear regression techniques are directly applicable to robotic systems such as pendulums but are not suitable for the ankle model identification. However, direct identification of the triple pendulum proved difficult in practice, so the shank model is identified separately from the thigh-torso model, which are combined to form the complete triple pendulum model. Each leg of the orthosis-dummy system is represented as two independent but identical triple pendulums, depicted in Fig. 2, when the foot is in complete contact with the ground and bilateral interaction at the hip and shared torso is neglected. The nominal dynamics of the system are u Yn n \u23a1 \u23a3 gc1 \u03c9\u03071 Y13 \u03c9\u03072 Y15 \u03c9\u03073 0 0 Y23 \u03c9\u03072 Y25 \u03c9\u03073 0 0 0 0 Y35 \u03c9\u03073 \u23a4 \u23a6 n (1) where Y13 \u2212L1s2q\u0307 2 2 + gc12 + 2L1q\u03081c2 + L1q\u03082c2 \u2212 2L1s2q\u03071q\u03072 Y23 L1s2q\u0307 2 1 + gc12 + L1c2q\u03081 Y35 s3 ( c2 ( L1q\u0307 2 1 \u2212 gs1 )\u2212 s2(L1q\u03081 + gc1) + L2(q\u03071 + q\u03072) 2) + c3 ( c2(L1q\u03081 + gc1) + s2 ( L1q\u0307 2 1 \u2212 gs1 ) + L2(q\u03081 + q\u03082) ) Y25 Y35 + L2c3\u03c9\u03073 \u2212 L2s3\u03c9 2 3 Y15 Y25 + L1c23\u03c9\u03073 \u2212 L1s23\u03c9 2 3 c j cos ( q j ) , c jk cos ( q j + qk ) , c123 cos(q1 + q2 + q3) s j sin ( q j ) , s jk sin ( q j + qk ) , s123 sin(q1 + q2 + q3) \u03c9k k\u2211 j 1 q\u0307 j \u03c9\u0307k k\u2211 j 1 q\u0308 j n \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 m1x1 + (m2 + m3)L1 I1 + m1x21 + (m2 + m3)L2 1 m2x2 + m3L2 I2 + m2x22 + m3L2 2 m3x3 I3 + m3x23 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 and the unknown inertial parameters are described in Table 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure23.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure23.2-1.png", + "caption": "Fig. 23.2 Geometrical dimension of a tube b coil, and c rod (dimensions are all in mm)", + "texts": [ + " The resultant Lorentz force which is a function of current density and magnetic field density also calculated in the same approach. Differential forms of Maxwell\u2019s equations (Eqs. 23.3\u201323.6) are shown below. r B \u00bc l0 J\u00fe e0@E=@t\u00f0 \u00de \u00f023:3\u00de r E \u00bc @B=@t \u00f023:4\u00de r E \u00bc q=e0 \u00f023:5\u00de r B \u00bc 0 \u00f023:6\u00de where E electric flux intensity, B magnetic field density, l0 magnetic permeability, J current density, q net charge density, e0 permittivity of vacuum, and t is for time. Geometrical dimensions tube, copper coil, and steel profiled rod are shown in Fig. 23.2a\u2013c. Total of four longitudinal grooves was designed, and its geometrical parameters kept constant for all simulation. Axisymmetric coil with dimension and shape as shown in Fig. 23.2b was used. Table 23.1 shows the properties of the aluminum tube, copper oil, and steel rod used in this study. Simulations were carried out for different voltages levels using identical coil geometry which is modeled as a rigid body to make it simple for calculation in the boundary element method. Hence, deformation was analyzed only on the tube at the joining zone. The ranges of voltage were varied from 1.5 to 4 kV with an increasing rate of 0.5 kV. The arrangement of workpieces and tool coil for simulation is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003864_ies50839.2020.9231879-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003864_ies50839.2020.9231879-Figure3-1.png", + "caption": "Figure. 3. (a) Diagram of Linier inverted pendulum model (LIPM) [1] (b) Movement pattern x-direction on uneven surface", + "texts": [ + " To find out whether the robot's posture has been balanced or not, a center of pressure (CoP) reading is done on the sole of the robot's feet. Corrections are then made if the CoP is not in its proper position. In addition to knowing the equilibrium of the robot's posture, CoP can also be used as a parameter to determine uneven footing. When the robot is on an uneven pedestal surface, CoP can be used to detect the level of the tilt of the pedestal. Henceforth, it is used to change the body part according to the slope, so that the robot will be able to keep running stable on the uneven surface. Figure 3 is an illustration of the linear control of the inverted pendulum model when the robot is running. This modeling will produce acceleration v. Inverted pendulum in the vertical direction will produce angular acceleration 0g = (,g/L) sin 0, while in the horizontal direction will produce angular acceleration 6Z = \u2014(v/L) cos 6. L in the equation is obtained from the equation L = y/l2 + Xa2. Total angular acceleration as follows : 0 = 0g + 6Z (1) From the equations that have been obtained before, the transfer function can be produced as follows : G(s) = \u2014 = = -------^2----- (2)v J v (s ) Ls2- g (tl s + 1 ) ( tl s +1) v j 214 Authorized licensed use limited to: University College London. Downloaded on November 02,2020 at 04:47:10 UTC from IEEE Xplore. Restrictions apply. In the equation 8, tl is time constant, which is obtained from the following equation: From the modeling in Figure 3 and the transfer function, the pendulum angle is obtained and the pendulum angle is used to control the balanced of the robot. The walking pattern is divided into two patterns, namely x-direction trajectory and z-direction trajectory. Figure 4 shows the trajectory robot in the x-direction(sagittal view). In the trajectory x-direction robot has two equations, xa for ankle trajectories dan xp for pelvis trajectories. xa(t) = (L0 + \u00bfi) - \u00bf s in (2ir^ ) ] cosp - L0 (4) xp(t) = Tii=oai ( 5 ) Figure 5 shows the ankle trajectory in the z-direction (coronal view) which has two trajectory steps" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002180_s00006-020-1045-1-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002180_s00006-020-1045-1-Figure2-1.png", + "caption": "Figure 2. Screw motion about the line axis l (ts: longitudinal displacement by d and Rs: rotation angle \u03b8): a motor relating two axis lines, b motor applied to an object, c degenerated motor relating two coplanar rotors. (Note: indicated 3D vectors are represented as bivectors in text.)", + "texts": [ + " But we must first show the relationship between motors and screw motion theory. Note also that the dual of a scalar is the pseudoscalar I and that the duals of the first three basis bivectors are actually the next three bivectors, that is, (e2e3)\u2217 = e2e3I = e4e1. According to Clifford [9], a motor operation is necessary to convert the rotation axis of a rotor into the rotation axis of a second rotor. Each rotor can be geometrically represented as a rotation plane with the rotation axis normal to this plane. Figure 2a depicts a motor action in detail. Note that the involved rotor axes are represented as line axes. In the figure, we first orient one axis parallel to the other by applying the rotor Rs of Eq. (16). Then, we slide the rotated axis a distance d along the connecting axis, so that it ends up overlapping the axis of the second rotor. Altogether, this operation can be described as forming a twist about a screw with line axis l, whose pitch relationship pitch equals d \u03b8 for \u03b8 = 0. A motor, then, is specified only by its direction and the position of the screw-axis line, twist angular magnitude, and pitch. Figure 2b shows an action of a motor on a real object. In this case, the motor relates the rotation-axis line of the initial position of the object to the rotation-axis line of its final position. Note that in both figures the angle and sliding distance indicate how rigid displacement takes place around and along a screw-axis line l, respectively. A degenerated motor is defined as a motor which can only rotate and not slide along the line l as Fig. 2c shows. In this case, therefore, the two axes are coplanar. 2.5. Motors, Rotors, and Translators in G+ 3,0,1 Since a rigid motion consists of the rotation and translation transformations, it should be possible to split a motor multiplicatively in terms of these two transformations, which we will call a rotor and a translator. In the following discussion, we will denote all bivector components of a spinor by slant bold lowercase letters. Let us now express this procedure algebraically. First of all, let us consider Eq", + " If we want to express the motor using only rotors in a dual spinor representation, we proceed as follows: M = RsT s = Rs(1 + I ts 2 ) = Rs + IRs ts 2 . (25) Let us consider carefully the resultant dual part of the motor. This is the geometric product of the bivector tS and the rotor Rs. Since both are expressed in terms of the same bivector basis, their geometric product will be also expressed in this basis, which can be considered as a new rotor R\u2032 s. Thus, we can further write M = Rs + IRs ts 2 = Rs + IR\u2032 s. (26) In this equation the line axes of the rotors are skewed, see Fig. 2a. That means that they represent the general case of non-coplanar rotors. If the sliding distance tS is zero, then the motor will degenerate to a rotor M = T sRs = ( 1 + I ts 2 ) Rs = ( 1 + I 0 2 ) Rs = Rs. (27) In this case, that is, when the two generating axis lines of the motor are coplanar, we get the so-called degenerated motor, see Fig. 2c. Finally, the bivector ts can be expressed in terms of the rotors using previous results \u02dcRsR \u2032 s = \u02dcRs ( Rs ts 2 ) ; (28) therefore, ts = 2\u02dcRsR \u2032 s. (29) Figure 2 shows that the 3D vector t, expressed in the bivector basis, is referred to the rotation axis of the rotor, and that tS is a bivector along the motor-axis line. Thus, t, considered here as a bivector, can be computed in terms of the bivectors tc and tS , as follows: t = t\u22a5 + t\u2016 = (tc \u2212 Rstc \u02dcRs) + (t \u00b7 n)n = (tc \u2212 Rstc \u02dcRs) + dn = tc \u2212 Rstc \u02dcRs + tS = tc \u2212 Rstc \u02dcRs + 2\u02dcRsR \u2032 s. (30) So far, we have analyzed the motor from a geometrical point of view. Next, we will look at the motor\u2019s relevant algebraic properties" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002299_rteict42901.2018.9012428-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002299_rteict42901.2018.9012428-Figure4-1.png", + "caption": "Fig. 4. Magnetic Parameters of Healthy LSPMSM. (a) Radial Flux Density . (b) 2-D Flux Density distribution. (c) Spatial FFT of Radial Flux Density distribution", + "texts": [ + " The electric, mechanical and magnetic parameters of this machine are recorded and analyzed. The parameters of a healthy LSPMSM are essential to be analyzed in-order to identify the expanse of deviation caused in these parameters when a fault is induced in the same LSPMSM. Electric parameters such as stator current and induced voltage as shown in Fig. 2, mechanical speed and torque parameters as shown in Fig. 3, and magnetic parameters such as Radial Flux Density and field lines distribution as shown in Fig. 4 are obtained from Maxwell. Stator turn-to-turn short-circuit fault is the most common electrical fault. This kind of faults occurs due to winding insulation problems, which are in-turn caused due to high temperatures and over-loading. Armature insulation can fail (a) due to the following reasons [5]: 1) Elevated stator winding temperature. 2) Slacken core lamination, joints and wedges. 3) Loose end- winding bracing. 4) Impurities such as oil, moisture and dirt. 5) Stresses while starting. 6) Electrical discharges" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003617_is48319.2020.9200119-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003617_is48319.2020.9200119-Figure2-1.png", + "caption": "Figure 2. Components of the IEC 63047 demonstration device.", + "texts": [ + " One or more mobile demonstration devices operate in a wireless network and send data to a control centre. The device can be built for less than 3700 \u20ac, making it ideal to be used in training and education, but also for developing sensor networks which could for example be deployed in case of CBRNE incidents. The relatively low cost and weight makes it also suitable for robotics applications. The device consists of a single-board computer extended with a GNSS receiver board and an antenna (see Figure 2). A radiation detection module is connected to the single-board computer via USB. Power is delivered from a battery pack via a switching regulator. Alternatively, a power bank may be used. The radiation detector used in the demonstration device is a Hamamatsu C12137 capable of sending the energies of the observed photons to the single-board computer at a rate of ten times per second. The clock of the single-board computer is synchronised with Universal Coordinated Time via the GNSS connection. The data acquisition software combines the radiation data with the position and encodes the data into the IEC 63047 format, using the open-source software asn1c developed by Lev Walkin [20]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000387_978-981-10-4938-5_13-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000387_978-981-10-4938-5_13-Figure6-1.png", + "caption": "Fig. 6 The principle of laser cladding", + "texts": [ + " 252 Powder Consolidation-Based AM Methods 253 Two primary machine tool types exist for the additive manufacture from metal 254 powders. They differ by design by one process employing a powder bed in which 255 powder is consolidated by either sintering or melting and where the process 256 can occur either in a vacuum or in an inert atmosphere. The other method employs 257 a cladding technique where a freely movable laser nozzle delivers a powder spray in 258 a carrier gas to the workpiece. 259 Cladding 260 The cladding principle is illustrated in Fig. 6. A tooling system comprised of an 261 integrated two-stage nozzle system function to deliver a steady stream of fluidized 262 powder to the workpiece while also delivering the heat flux in the form of focused 263 laser light to the workpiece. Fluidizing the powder and bringing it into the beam of 264 the laser are done by means of an inert carrier gas spraying the powder directly to 265 where material is required. The laser beam intercepts the powder stream at the 266 desired deposition location and fuses the melt directly onto the workpiece" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001440_icca.2019.8899716-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001440_icca.2019.8899716-Figure1-1.png", + "caption": "Fig. 1. Kinematic structure of a five-axis engraving machine.", + "texts": [ + " Synchronization of tool tip and rotary axes are guaranteed while traversing along tool paths. A bi-directional scanning (BS) algorithm is utilized to plan the overall feedrate profile. Finally, simulations and experiments are carried out on a desktop five-axis engraving machine to validate real-time performance, contour accuracy and cycle time compared with a commercial Heidenhain controller. A desktop five-axis engraving machine with a swivel head and a rotary table is used in this paper. The kinematic structure of the five-axis machine is shown in Fig. 1. The transformation matrix wTt which describes the tool tip with respect to the workpiece frame can be derived, and the forward and kinematic equations (1) and (2) are obtained bzcbycbx wbbtbzz btcbycxcy btcbycxcx cossocso ZZcp Zsscsp Zcsscp =\u2212== +\u2212= ++\u2212= \u2212+= ,, \u03b8 \u03b8\u03b8 \u03b8\u03b8 (1) )arctan(),arccos( xyczb wbbtbzz ycxcy btbycxcx ooo ZZcp pcps Zspspc \u2212== \u2212+= += +\u2212= \u03b8\u03b8 \u03b8 \u03b8 \u03b8 (2) where five-axis motion commands \u03b8m, tool tip position Pwt and tool axis orientation Owt are defined as [\u03b8x \u03b8y \u03b8z \u03b8b \u03b8c]T, [px py pz]T and [ox oy oz]T, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002614_j.promfg.2020.04.235-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002614_j.promfg.2020.04.235-Figure2-1.png", + "caption": "Fig. 2. (a) Base plate with welding tracks, (b) disc made of a rod with the used upsetting specimen", + "texts": [ + " Uniaxial compression tests were carried out to determine flow curves of the investigated materials 41Cr4 and X45CrSi9-3. In this paper, the flow behavior of two specimens extracted form deposition material and conventional material was investigated. At first, cylindrical samples were taken from a welding track by wire eroding. The welding track was welded on a base material C22.8. The specimens for the second variant of sample preparation were taken from conventional rod material. The rod has a diameter of 40 mm. Fig. 2 (a) shows the baseplate and the welding track from the first preparation method and in Fig. 2 (b) the rod as well as the used specimen within this study are depicted. The specimens had a diameter of 5 mm and a height of 8 mm. The upsetting tests were carried out with the forming and quenching dilatometer DIL 850A/D+T from TA instruments. For the tests three different temperatures (900 \u00b0C, 1050 \u00b0C, 1200 \u00b0C) and three different strain rates (0.1 s-1, 1 s-1 and 10 s-1) were investigated. The specimens were inductively heated. The samples were upset to a plastic strain of \ud835\udf11\ud835\udf11 = 0.7. Up to this point, the influence of friction on the measurement result is negligibly small [11,12]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002029_iccas47443.2019.8971518-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002029_iccas47443.2019.8971518-Figure9-1.png", + "caption": "Fig. 9 Numerical results on expandable drill bit.", + "texts": [ + " Simulation is performed under the condition that the drill bit module rotates at 500 rpm. Pressure and rotation are applied to part 3 (ground) to implement rotational motion and pressure, and part 1 and part 4 are fixed. Part 1, 3, and 4 set as a rigid material model and part 2 is configured as a plastic kinematic material model. All elements of the model are made up as solid elements. Part 1 and 2 / part 2 and 3 are set up to be able to make reaction force with each other using surface to surface contact algorithm. The simulation result is illustrated in figure 9. From the numerical analysis, it is confirmed that the area where the stresses are concentrated at the expandable drill bit. As shown in figure 9, the stress is concentrated on the section where the gear and blade are connected. The part with the greatest stress is subjected to the force of about 10 MPa. However, the durability problem of the designed system is not expected through the material properties of the steel when drilling 3 MPa concrete. In this paper, the excavation mechanism of a small-sized embedded robot for shallow drilling is proposed, and the design and analysis are conducted. A hybrid type drilling mechanism is designed by mimicking the incisors of the naked mole-rat and the forelimbs of the European mole" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000297_kem.799.263-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000297_kem.799.263-Figure9-1.png", + "caption": "Fig. 9 Modified LCD design Fig. 8 Improved lift-swing motion mechanism", + "texts": [ + " The space between the wall of the nozzle support and the sliding component (KLP-P-007 modif.) was reduced and the spindle (KLPP-008 modif.) was attached to a wall with a fixation plate (KLP-P-16). This modification offered significant compression stress reduction, from 1.547 MPa to 0.413 MPa proving that the weight distribution got more even. Last, all the components shown on the right side of Fig. 8 were submitted to tolerance adjustment redesign, in accordance with ISO DIN 2768-CLASE cL. With all the modifications, the improved LCD look like the right-hand diagram in Fig. 9. As shown in Fig. 9, apart from the self-aligning bearing in the swing-lift motion area, shown in Fig. 8, additional bearings were included in the design to improve the transversal movement. They are marked with a red circle in the right image. To summarise the improvements, the main technical characteristics are shown in the table below. As can be seen in Table 4, the mobility of the device has increased, and its weight has been reduced by two kilos. Mechanically speaking therefore, the device has undergone significant improvements, using finite element analysis (FEA) methods and CAD tools" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure11-1.png", + "caption": "Figure 11. Eight-bar mechanism.", + "texts": [], + "surrounding_texts": [ + "The primary objective of this study was to identify the joint that producesthe highest error in a mechanism due to clearance. Simulations were carried out for all mechanisms described in the previous section by introducing clearances in different joints, one at a time. The effect of different speeds (of crank) and clearance sizes was also studied. Some important results are discussed here. The results for the four bar mechanisms have been summarized in figures 13, 14 and 15. These show that Joint 3 is the most critical joint and the values vary significantly for 0.02 mm but they are very similar for 0.5 mm. So, sensitivity to speed is high for low clearance. Figures 16, 17 and 18 show mean deviation at different joints of different inversions of the Watt chain at 0.1 mm radial clearance. The trend of all the curves is the same. Joint 3 is the most critical joint in all the three mechanisms followed by joints 2, 4 and 5, i.e., the criticality order here is: Joint 3[ Joint 2[Joint 4[ Joint 5. Joint 6 in figures 5 and 6 is the grounded joint which is least critical of them all. Figures 19 and 20 show the criticality graphs for Stephenson1 mechanism and Stephenson1 mechanism with slider, respectively. Both the graphs show the same trend where the deviation is maximum at the output link and decreases as one proceeds to the other end of the mechanism. A common trend which seems to emerge is that the first joint next to the crank, which has a link connected to ground, is the most critical joint. After that the criticality ranking goes towards the crank and then goes further down to the joints in order of their distances from the crank. We would examine the validity of this hypothesis in the rest of the study. Figures 21 and 22 show the criticality graphs for Stephenson2 mechanism (1) and Stephenson2 mechanism (2). The order of criticality for both the mechanisms is: Joint4 [ Joint3 [ Joint2 [ Joint5. This is as per the hypothesis stated earlier. The most critical joint is the joint next to the crank, which has a link connected to the ground (i.e., Joint 4). The next most critical joints are the ones towards the crank, i.e., Joint 3 and Joint 2. Then comes the joints which are left, i.e., Joint 5. Joint 6 is the grounded joint. The two graphs look very similar as the structures of mechanisms are the same. The deviations increase significantly with increase in speed, if the clearance is at Joint 4. The values in first graph are slightly higher than the corresponding values in the second graph. The overall conclusion is that deviations are more, if the output link is closer to the crank. Similar results are seen in eight and ten-bar mechanisms as shown in figures 23 and 24. The joint next to the crank is the most critical joint. It should also be noted that the grounded joints also show a trend. The grounded joint in the first loop has a significant effect but rest of the grounded joints have very little effect on the output. Error in output generally increases with increase in the crank speed (expect for an aberration shown in figure 19). This happens for all clearance sizes. However, the change in mean deviation is quite small even with a 5 times increase in speed from 100 to 500 rpm. Figures 25 and 26 show the variation of mean deviation with increase in speed when a clearance of 0.1 mm is at Joint 2 and Joint 3, respectively. It can be noticed that the deviations usually increase with speed. This means that a machine working at higher speeds is expected to produce more error than a machine working at lower speeds. Increase in radial clearance size brings an almost linear increase in the deviation values as seen in figures 27 and 28. It is also worth noting from figure 27 here that the deviation of slider is nearly 50% for 0.1 mm clearance (0.0501 is 50% of 0.1) and it goes on increasing to 80% for 0.5 mm clearance (0.3977 is 80% of 0.5). This happens because at lower values of clearance, the number of collisions between journal and bearing is more. Due to more number of collisions, the journal is sent back to the center again and again, and thus it remains at the center for a larger amount of time as compared to the case of large clearances. Thus, in larger clearances, the journal appears to remain closer to the walls of the bearing for an extended period and spends less time going to the diametrically opposite wall." + ] + }, + { + "image_filename": "designv11_80_0000280_s10999-019-09455-z-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000280_s10999-019-09455-z-Figure1-1.png", + "caption": "Fig. 1 Typical epitrochoidal star-ring set and flow distributor valve of an ORBIT motor", + "texts": [ + " The unmodified epitrochoidal profile (two lobes) is used to shape the outer fixed member of a Wankel engine. The envelope (three lobes) of the epitrochoid is used as the inner rotary member. In case of HST units the constant difference inward shifted epitrochoid is used to shape the inner member (star). The envelope lobes become circular at active contacts. It is used as outer member (ring). Such a modification of epitrochoid is done to have the better gearing action. The envelope lobes may be integral with the ring or rollers placed within the partial bush in the ring (Fig. 1). Such ROPIMA type HST units, being simple in construction and of low costs, possess many positive features. However, their use is still within medium pressure application range due to poor leakage characteristics. The leakages are mainly through the flow distributor valves, the side plates and through the generated gaps at transition active contacts (inter-chamber leakages). Improving the leakage characteristics of this class of HST units is still a challenging area of research. Volumes trapped in chambers formed by higher pair geometrically form-closed contacts (named as \u2018active contacts\u2019 in the present investigation) between the star and ring lobes (teeth), undergo volume changes. This is due to contraction and expansion of the in plane form-closed areas between two consecutive active contacts during rotations. This causes pumping and motoring actions in such machines. Among the active contacts, two contacts separate low pressure zone (LPZ) chambers from the adjacent high pressure zone (HPZ) chambers (Fig. 1a, b). These critical contacts are called as \u2018transition\u2019 contacts hereinafter. The models shown in Fig. 1 are of said constant difference modified epitrochoidal star with its envelope ring. The fixed axes (GEROTOR) units, where both the star and the ring rotate about their respective central axis (as in fixed axes internal gearing) (Fig. 1a), possess high speed low torque (HSLT) features. In such units, lobes are mostly made integral with the ring as shown in Fig. 1a. In the other kinematic version, usually the ring is held fixed and the star is allowed to rotate in an epicyclic/orbital manner. The rotation of the star about its own axis and the torque are transmitted through a cardan shaft. It is to be noted that the shaft can be held fixed and output motion is taken through the outer ring. For this some constructional features like inlet\u2013 outlet valves are to be altered. However, such units are used only as low speed high torque (LSHT) motors. They are popularly known as ORBIT motors. Figure 1b shows the star-ring set (with separate rollers as lobes) commonly used in an ORBIT motors. Such a set with roller lobe ring is sometimes called as GEROLOR. Figure 1c shows the inner view of an ORBIT motor with pintle type flow distributor valve. This valve is integral with the fixed axis output shaft. The ring is bolted to the motor housing and remains fixed. The rotation of this valve is to be same as the rotation of the rotor about its own axis. The rotational motion from rotor to the valve and output shaft is connected through a cardan shaft, as the rotor rotates in epicyclic manner. Opposite to the output end, the shaft is hollow with splines at the inner end. The star centre is also splined. The cardan shaft has crowned gear teeth at both ends. The kinematic behavior of a cardan shaft is same as that of a tail shaft, with universal joints, in automobiles. The angular positions of all splines, lobes in star and flow distributor grooves on the valve are in sequence for the function of the machine and are manufactured carefully. The valve sequences are discussed in detail by Maiti (1990, 1992). In ORBIT motors (Fig. 1c), pressurized fluid is forced into a chamber in expanding mode, causes the inner member to rotate about its own axis and to revolve around the centre of the stator i.e., the ring. After a certain angle of rotation (depending on the number of chambers i.e., lobes in ring) the adjacent chamber, which was in compression mode, begins to expand due to flow-in of pressurized fluid. The expanding chamber enters into compression mode just after reaching its maximum volume. The passage through the flow distributor valve (Roy et al", + " Moreover, for gearing action sealing has to depend only on metal to metal contact (Maiti 1990). The form-closeness at two \u2018transition\u2019 contacts in all such units is lost during the operation. The contacts are separated due to the occurrence of contact deformations at other load transmitting active contacts, for a period of time during each power cycle (Maiti 1990, 1991, 1993). In a GEROTOR unit the kidney port valve, which remains fixed, is used as inlet\u2013outlet valve. Usually the star is coupled to the drive shaft with well designed key or spline (Fig. 1a). The ring rotates inside a bush under hydrostatic lubrication environments. The design is made such that no or negligibly small radial force is transmitted to the drive shaft. This is to ensure better geometric form-closed contacts of star-ring at their active contacts as well as better lubrication effects for ring rotation. Overall, this would improve the GEROTOR pump and motor performances. With such GEROTOR HST unit assemblies, the star-ring behavior would be more or less same as in ORBIT motor unit with proper kinematic correlation (Maiti and Sinha 1988\u20132013, 1990; Maiti 1991, 1993)", + " In the present investigation we have established an FEM approach to estimate the gaps at transition contacts in confirmation toMaiti\u2019s TEM approach.We also estimated the inter chamber leakages through the gaps at transition contacts. 2.1 Expansion-compression chamber actions We have considered the ORBIT motor for the present investigation. The inner gear (star) of the motor is the rotor. It has epicyclic gear motion inside the envelope lobed (circular roller) fixed ring gear (stator). Henceforth the terms \u2018star\u2019 as well as \u2018rotor\u2019 are used for inner member, and the terms \u2018ring\u2019 as well as \u2018stator\u2019 are used for the fixed outer member (Fig. 1b) in the present report. In the shown illustrations the ring has Z (= 7) lobes (teeth) and the start has one less (Z - 1) lobes (teeth). The number of chambers is 7. It is to be noted that in such epitrochoidal generated set the number of chambers is always equal to number of lobes of the ring. It is to be noted that odd Z give low flow/torque ripple in comparison to that with even Z of higher integer. The motor experiences Z(Z - 1) number of total chamber or piston actions (expansion-compression) in one full revolution of the output shaft i", + " The ring segment between two bolts is considered as beam and it is allowed to deflect elastically in radial direction. The boundary conditions are assigned accordingly. FEM results showed large deformations of the ring. These results are also unrealistic. Type C\u2014the final model: There must have some other forces acting on the ring that holds the ring, allowing its elastic behavior, as revealed from the analyses in two earlier models. We now consider the side thrusts on the ring from both the valve plate and the end plate (Fig. 1c). This is due to bolt tightening forces for fixing the ring with the motor body. These forces are to be sufficient to transmit the reaction torque from ring to the motor body as well as to set the seal for preventing external leakages. In practice each bolt is tightened for a specific torque by a torque wrench. We consider a distributed thrust load on the annular ring due to the bolt tightening forces. We assume that the distribution pattern, on the span between two bolts, is same all over the ring", + " circular arc) of radius rm, is spread over twice the leaning angle, /. The first crunode is at w \u00bc p=Z with respect to X0CY0 axis system. Other crunodes are at 2p=Z angle from each other. The circular arcs of useful spread of 2/ angle or slightly more can be attached to all crunodes and then these arcs are joined with suitable identical arcs covering the envelope to make integral profile of the ring. Else in a ring plate with partial bush cavities at all crunodes, housing rollers of radius rm and length equal to the width of the ring b, is used (Fig. 1). The position vector of the modified epitrochoid generating point M, with respect to the centre of star O can be expressed as: OM ! \u00bc C~0 \u00fe A~0 \u00fe r~m \u00f012\u00de As already mentioned for unicursal epitrochoid and its envelope the following condition has to be satisfied. R0 r0 \u00bc Z Z 1\u00f0 \u00de i.e., R0 R0 r0\u00f0 \u00de \u00bc Z \u00f013\u00de The coordinates of point M (Xm, Ym) of the modified epitrochoid in dimensionless form, with respect to XOY co-ordinate system, are as follows: Xm \u00bc A0 cosw\u00fe 1 Z cos Zw rm cos w\u00fe /\u00f0 \u00de Ym \u00bc A0 sinw\u00fe 1 Z sin Zw rm sin w\u00fe /\u00f0 \u00de \u00f014\u00de The leaning angle (/) is an important parameter in the modification of epitrochoid profiles because in gearing action (p=2 /) may be considered as the instantaneous pressure angle. In dimensionless parametric form it is expressed as: / = tan 1 sin Z 1\u00f0 \u00dew A0 \u00fe cos Z 1\u00f0 \u00dew \u00f015\u00de where w \u00bc Zn \u00f016\u00de Figure 9 shows a typical epitrochoid generated star-ring (with roller) set. Appendix 2 Calculation of side thrust on the ring due to tightening it between the end plate and valve plate by fixing bolts (Fig. 1c) for Type-C FEM model Each side flat area of the ring of this considered ORBIT motor unit is 3143.17 mm2, as estimated by solid modelling. Each bolt (M8 9 1.25) cross sectional (effective) area As, is 36.6 mm2. Number of bolts 7. We assume the Proof Strength (ry) as 1098.34 MPa considering the steel\u2019s strength class 12.9. Co-efficient of friction (fn) for metal to metal (steel) contact is 0.4. Assuming that 70% of proof strength is applied on each bolt tightening force, axial force experienced by each bolt is 0:7 ry As N \u00bc 0:7 1098:34 36:6 N \u00bc 28139:47 N" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002904_phm-besancon49106.2020.00042-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002904_phm-besancon49106.2020.00042-Figure3-1.png", + "caption": "Fig. 3. Tightening Tool.", + "texts": [ + " These two modules scale the input and output of the fuzzy controller, and are the input and output interfaces of the fuzzy controller. They not only make the front and back modules match, but also improve the performance of the fuzzy controller. The functions of the 3 modules in the core of fuzzy controller are: fuzzification module D F converts digital quantity into fuzzy quantity, A* R operates approximate reasoning based on the input fuzzy quantities to obtain a fuzzy quantity U, defuzzification module F D converts fuzzy quantity into digital quantity. A common tightening tool is shown in Fig. 3. It contains a servo motor(including an encoder), a reducer, a torque sensor and a screwdriver. We use the torque sensor to obtain the torque value during the assembly process, and then design the controller to control the motor speed, that is, the rotation angle of the bolt. Then according to the relationship between the pretightening force and the rotation angle shown in Fig. 1, we can Authorized licensed use limited to: University of New South Wales. Downloaded on July 26,2020 at 23:21:41 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002791_s11431-020-1569-6-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002791_s11431-020-1569-6-Figure2-1.png", + "caption": "Figure 2 Mechanical analysis of an SLE pantographic mechanism under axial loads.", + "texts": [ + " It should be highlighted that the smaller of P1 and P2 values is chosen as the critical load of the SLE. When the mechanical analysis and stability analysis of an SLE are completed, the next step is to focus on the buckling conditions of the deployable structure according to the contribution of each SLE. Deployable structure consists of units which are composed of pairs of bars connected at a joint that allows a compact and deployable configuration. Many identical SLEs are arrayed along x axis to form a linear array structure, that is, based-SLE deployable structure [44] shown in Figure 2. In the SLE pantographic mechanism, this structure is assembled from right to left, and the last unit is hinged on the fixed surface. At the same time, all units will theoretically exist in one plane, and the two nodes in the first SLE will bear horizontal axial loads, P. In order to study the effect of the load on the structural buckling, each unit is analyzed in accordance with the assembly direction, and the analysis results are shown in Figure 2. In the global coordinate system o-xy, each unit is numbered from right to left, and the number of elements and deployed angle are n and \u03b3, respectively. For simplification, use A, B, C, D and O to indicate the hinge point of each unit, there will be no ambiguity because each unit has a subscript. In addition, due to the linear superposition of the axial loads, the forces that bars underwent at the hinges varies linearly from right to left. Also, we ignore the friction effect at the hinge and taking i-th unit for analysis, its force at the joint can be expressed as F P F iP F P F i P F P F iP F P F i P = , =2 tan , = , = 2( 1) tan , = , = 2 tan , = , = 2( 1) tan , (12) Ax i A i Bx i By i Cx i Cy i Dx i Dy i ( ) y ( ) ( ) ( ) ( ) ( ) ( ) ( ) where \u03b3 is the deployed angle, and \u201c(i)\u201d represents the i-th unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003450_s10958-020-05007-5-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003450_s10958-020-05007-5-Figure1-1.png", + "caption": "Figure 1. The boundary of the parametric resonance domain (3.7) in the space of parameters h, \u03b3, \u03b4 (a) and the cross-section of the domain by the plane \u03b4 = const (b). The third order resonance curve is denoted by \u03b33.", + "texts": [ + "6) The equation for the boundary of the parametric resonance domain in the original parameters is looked for in the form of series \u03b3(\u03b4, h) = \u03b30(\u03b4)+\u03b31(\u03b4) \u221a h+O(h), where \u03b30(\u03b4)\u00b1 \u221a\u22124\u03b43+17\u03b42\u22124\u03b4 5\u03b4 determines the dependence of \u03b3 on \u03b4 for the exact resonance \u03c92 = 2\u03c91. Substituting this series into (3.6) and equating coefficients of like powers of h, we find \u03b31(\u03b4) and obtain the equation for the boundary of the parametric resonance domain in the space of parameters h, \u03b3, \u03b4: \u03b3(\u03b4, h) = \u03b30(\u03b4)\u00b1 |3\u03b4 \u2212 2|\u221a6 4 \u221a 5 + 5\u03b4 25 \u221a \u03b4(4\u2212 \u03b4) \u221a h+O(h). (3.7) The surface defined by (3.7) is shown in Figure 1. For the values of \u03b3, \u03b4 close to the resonance ones and small h this surface well approximates the boundary of the parametric resonance domain which is confirmed by the fact that the results of analytic and numerical analysis coincide. To find the family of long-period motions, we use the methods described in [3]. Applying a transformation x1, x2, y, y2\u2192Q1, Q2, P1, P2, we reduce the Hamiltonian (3.2) to the form KIII = P1 + \u03bcP2 + \u03b5(P1 + 2P2) \u221a P2 sinQ2 +O(\u03b52), (3.8) where P1 2P2. The system with the Hamiltonian (3", + " This system is reduced at level of the first integral P1 = C by the approximate system whose Hamiltonian function is the normalized part of the Hamiltonian (3.8). The equilibrium states Q\u2217 2, P \u2217 2 of the system with the Hamiltonian (3.9) are determined by Q\u0307\u2217 2 = \u2202KIV \u2202P \u2217 2 = 0, P\u0307 \u2217 2 = \u2212\u2202KIV \u2202Q\u2217 2 = 0. (3.10) The number of equilibrium states depends on the parameter \u03bc. In the original variables \u03c8, \u03b8, p\u03c8, p\u03b8, these equilibrium states near the third order resonance correspond to two families, denoted by \u03931 and \u03932. of long-period motions. Figure 1 (b) shows the existence domains for these families in a neighborhood of the third order resonance curve. The motions of the family \u03931, exist in the domains I and II, and the motions of the family \u03932 exist in the domains II and III. For small \u03bc the boundaries of the domains I, II, III coincide with the boundary of the parametric resonance domain and are described by Equation (3.7). In a neighborhood of the resonance \u03c92 = 2\u03c91, the families \u03931 and \u03932 of long-period motions are orbitally stable in the linear approximation [3]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001959_s0036029519130111-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001959_s0036029519130111-Figure1-1.png", + "caption": "Fig. 1. Scheme of the position of the model in space.", + "texts": [ + " To solve the heat conduction problem, we used an axisymmetric two-dimensional model of the tool. The tool was inserted into the chuck by 25 mm. The chuck is made of steel DIN-C15 taken from the library of Deform and the tool is made of DIN-AlMg5 (Deform library). The elements are square, four-node, 1 \u00d7 1 mm in size. Heating was performed in 10 s (Fig. 4). The following boundary conditions are set to calculate the loss of stability during primary tool\u2013substrate contact (during contact and the first 5 seconds): a force of 15 kN is applied to the chuck from side +Z to \u2013Z (Fig. 1), the rate of chuck rotation about axis Z is 3.715 rad/s (35 rpm), the substrate and the chuck are absolutely rigid bodies, sticking contact at the chuck\u2013 tool boundary, and the contact at the tool\u2013substrate boundary is 0.4 according to Siebel. The temperature problem was not solved in parallel. The substrate made an angle of 2\u00b0 with the horizontal plane, which imitates the real tool inclination during coating deposition, which is used in practice. When the tool is in contact with the substrate, the intensity of internal stresses increases to the ultimate tensile strength of the material at room temperature (300\u2013310 MPa) at the lower edge under the action of an axial load and friction at the tool\u2013substrate boundary (Fig", + " To calculate the deformation of the consumable tool during coating deposition, we superimposed the temperature fields obtained at the first stage of calculation on the developed models. Here, we solved the deformation problem and did not take into account the heat losses via convective exchange with the environment and the heat exchange between the tool and the substrate. We also did not specify a temperature boundary condition for heating the tool. As in the previous case, a force of 15 kN is applied to the chuck from side +Z toward \u2013Z (Fig. 1), the rate of chuck rotation about axis Z is 3.715 rad/s (35 rpm), and no translational motion is set to the tool in the +X direction. The calculated deformation time is 2 s and the time step is 0.05 s. The calculation shows that the internal stress intensity reaches 50\u2013100 MPa (the ultimate tensile strength of the material is 60\u201370 MPa [10]) in the tool\u2013substrate contact zone, where the tool material temperature is 480\u2013530\u00b0C. It should be noted that the zone of high (at given temperatures) internal stresses, which cause a plastic f low, is wider than that during cold upsetting of the rotating tool" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003949_5.0024901-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003949_5.0024901-Figure2-1.png", + "caption": "FIGURE 2. Staircase Dimensions FIGURE 3. Nose Line Climbing", + "texts": [], + "surrounding_texts": [ + "security setups. Hence, it is necessary to keep up with the evolving technology and countering to threats. Such robots can play a vital role in surveillance and gathering intelligence from places where deploying personnel could be risky. Hence, it is necessary to manufacture such a vehicle which would face no hindrances while traversing through varied terrains depending on the power output. Apart from the power output, the weight of the vehicle is taken into consideration as heavy weight vehicles will consume more power and less movement.\nIt is very necessary to maintain the balance of the CaterBOT while it traverses on rough terrain and hence manufacturing should be done with extreme care. After finishing with the design and calculation processes, the manufacturing of the chassis of the \u2018CaterBOT\u2019 commenced\nAluminium was used to initially considering the light weight. 12 V direct current motors were bought to power the driver wheels. The flipper movement was going to be achieved using a servo motor. The signals to the servo was remote-controlled and the actuation was limited a pulse. The flipper used to lock in its position according to the input. After installing the battery and the camera on the chassis, the first test run was made. The chassis failed as it could bear the load of the mounted electronics equipment. Neither was the motor able to supply enough power to the driver wheels nor was the battery providing enough voltage for smooth operations. Hence, it was necessary to change the material of the chassis. The focus shifted to acrylic which lighter than aluminium and has better load bearing. Instead of using servos for the flipper movement, worm motors were used to change the angle of the flippers according to the size and shape of the obstacle. Relay Cards connected to these motors were actuated digitally using a mobile phone application which commanded the vehicle to move. An Arduino controlled all the movements on this vehicle. After reinstalling the camera and the battery, a second run test was conducted. It yielded the required results as the \u2018CaterBOT\u2019 was able to climb over the obstacles.\nIt is very essential to maintain the balance of the vehicle while traversing.Aluminium was used to initially considering the light weight. 12 V direct current motors were bought to power the driver wheels. The flipper movement was going to be achieved using a servo motor. The signals to the servo was remote-controlled and the actuation was limited a pulse. The flipper used to lock in its position according to the input. After installing the battery and the camera on the chassis, the first test run was run. The chassis failed as it could bear the load of the mounted electronics equipment. Neither was the motor able to supply enough power to the driver wheels nor was the battery providing enough voltage for smooth operations. Hence, it was necessary to change the material of the chassis. The focus shifted to acrylic which lighter than aluminium and has better load bearing. Instead of using servos for the flipper movement, worm motors were used to change the angle of the flippersaccording to the size and shape of the obstacle. Relay Cards connected to these motors were actuated digitally using a mobile phone application which commanded the vehicle to move. An Arduino controlled all the movements on this vehicle. After reinstallingthe camera and the battery, a second run test was conducted. It yielded the required results as the \u2018CaterBOT\u2019 was able to climb over the obstacles.\n020015-2", + "The factors to calculate stair angle are hiking efficiency, availability of space in building and easiness to walk. The line which is connected the stair nose is known as nose line [7]. Stair angle which happens between horizontal ground and nose line. Stair angle usually changes from 20\u00ba to 45\u00ba. But then most suitable angle is 30\u00ba.\nThe length of foot, space of step and type of building are factors for selecting the stair dimensions. The stairs measurementcontains the flipper height and width of thread. The height of flipper usually varies from 0.12m to 0.18m by relates with uplift distance of human step. The width of thread usually varies from 0.2m to 0.35m by relates with human foot length. If the stairs dimension increases, it will be more difficult for the climbing robot [8] It is important to design the robot\u2019s dimension reasonably as the stairs dimension is fixed in a range. The process of stairs-climbing can be divided into three phases, that is, flipper climbing, flipper crossing, and nose line climbing. Both flipper climbing ability and flipper crossing ability also reflect the vertical obstacle climbing ability of the robot.\nThe CaterBOT\u2019s flipper climbing is same as that of the tank\u2019s vertical wall climbing mechanism. The process of flipper climbing is a dynamic process, but research of various journals has used statistical formulation to calculate the flipper climb of the secondary belt drive.\n020015-3", + "By the end of the flipper climbing process the robot will come have to clear the step nose. The robot should cross the flipper at a specific height. The maximum height of obstacle that it can cross is given by the equation.\nH = r + (l/2 + d) sin \u03b8 \u2013 r/cos\u03b8 Where \u2018H\u2019 being the height of the obstacle, \u2018r\u2019 being the radius of the wheel, \u2018l\u2019 being the center to center distance\nbetween the two wheels of the flipper and \u03b8 the angle of inclination of the robot.\nThe real problem arises when the flipper crosses a step and reaches the nose line of the step the following three conditions are to be satisfied to successfully complete this stage. Center of mass point condition comes up when the center of mass of the robot overcomes the nose of the step.\nl/2 + d \u2265 \u221a (b2 + h2) The contact point condition appears when the front contact point of the drive crosses the second nose, i.e. l \u2265 2 \u221a (b2 + h2) \u2013 r Tan \u03b8 The Friction condition states that the resultant force should be greater than the resolved gravity force\nalong the direction considered. \u2211 Frt = G sin \u03b8\nOur team came up with an idea of co-axial shafts but, yet the designing of shaft was a hectic process, on the more power transmission was another big problem to cope up with, which is discussed ahead in upcoming sub-headings. A co-axial shaft as the name suggests uses two shafts whose axis of rotation is same, so a co-axial shaft consists of a hollow shaft rotating in the outside while there\u2019s a solid shaft which has its individual rotational freedom while maintaining the same axis of rotation. There were two parts to design of the co-axial shaft that\u2019s is the design of hollow shaft and solid shaft both bearing different kinds of loads.\nThe power transmission mechanism is done via belt drive. Basic dimensions for the gear are tabulated below.\nThe motor provides 100 rpm of speed which in turn is converted into 33 rpm roughly.The worm motors generate a torque of 35 kgf which rotates the gears and therefore the shaft assembly connected to the inner wheels of the\n020015-4" + ] + }, + { + "image_filename": "designv11_80_0002878_0954406220930052-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002878_0954406220930052-Figure9-1.png", + "caption": "Figure 9. An eight-bar linkage illustration.", + "texts": [ + " To synthesize this type of eight-bar linkage, 1\u00fe n\u00fe n m SCs need be calculated, and there are 18 eight-bar linkages that belong to this type synthesis, as shown in Figure 24. In this paper, the circuit and branch defects are collectively called defects, because, the circuit and branch defects can coexist for a synthesized eight-bar linkage solution that cannot move smoothly through task positions. This section introduces the method and process to determine whether the linkages have defects based on the Jacobian matrix. The Jacobian matrix of the eight-bar linkage Using the notation of Figure 9, the closed-loop equations of the eight-bar linkage are labx \u00fe l5 cos 5 \u00fe l4 cos 4 l2 cos 2 l3 cos 3 \u00bc l1 cos 1 \u00f05\u00de laby \u00fe l5 sin 5 \u00fe l4 sin 4 l2 sin 2 l3 sin 3 \u00bc l1 sin 1 \u00f06\u00de labx \u00fe l5 cos 5 \u00fe l6 cos\u00f0 4 \u00fe \u00de l7 cos 6 \u00bc l8 cos\u00f0 1 \u00de \u00f07\u00de laby \u00fe l5 sin 5 \u00fe l6 sin\u00f0 4 \u00fe \u00de l7 sin 6 \u00bc l8 sin\u00f0 1 \u00de \u00f08\u00de labx \u00fe l5 cos 5 \u00fe l11 cos\u00f0 4 \u00fe \u00fe \u00de l10 cos 7 \u00bc l9 cos\u00f0 1 \u00de \u00f09\u00de laby \u00fe l5 sin 5 \u00fe l11 sin\u00f0 4 \u00fe \u00fe \u00de l10 sin 7 \u00bc l9 sin\u00f0 1 \u00de \u00f010\u00de where labx\u00bc b0x\u2013a0x, laby\u00bc b0y\u2013a0y, (a0x, a0y) denotes the coordinates of joint a0, (b0x, b0y) denotes the coordinates of joint b0, and h1 is the input angle", + " Equations (5) to (10) form a set of nonlinear simultaneous equations with six unknowns hj (j\u00bc 2. . .7). Differentiating both sides of equations (5) to (10) with respect to h1, the determinant of the Jacobian matrix is d \u00f0 1\u00de \u00bc 1 2l2l3l5l7l10 \u00f011\u00de where 1 \u00bc sin\u00f0 2 3\u00de 2 \u00bc sin\u00f0 7 5\u00de sin\u00f0 6 4 \u00del6 \u00fe sin\u00f0 4 \u00fe \u00fe 7\u00de sin\u00f0 6 5\u00del11. Generally, the sign of d(h1) ( 1 or 2) changes at limit positions of the driving link, i.e. at the singularity configuration. Note that the two-bar Assur group of this eight-bar linkage (Figure 9) has singularity configuration when 1\u00bc 0, and the four-bar Assur group has singularity configuration when 2\u00bc 0. Thus, the eightbar linkage singularity configuration occurs when any of the Assur groups for the eight-bar linkage has singularity configuration. So, if the sign of i changes, the eight-bar linkage is found to have defects. Similarity, the eight-bar linkage can also consist of three two-bar Assur groups (such as Figure 10(a)), or one six-bar Assur groups (as seen in Figure 10(b)). Defect identification for eight-bar linkages A linkage changes either circuit or branch when the sign of i changes. However, even if i keeps the same sign in the task positions, the eight-bar linkage may be still a defective linkage. For example, Table 1 lists the parameters of an eight-bar linkage (Figure 9) at two positions and 1, and 2 of d(h1). Figure 11 shows the change curves of 1 and 2 in the motion domain of the eightbar linkage. The values of 1 and 2 have the same sign in both positions. However, the eight-bar linkage has to change branches when it moves from position 1 to 2, since 2\u00bc 0, i.e. d(h1)\u00bc 0 occurs. As shown in Figure 11, if the eight-bar linkage needs to move from position 1 to 2, 2 changes along the green arrow, and the input angle cannot vary monotonically from \u2013 30 to 60. Therefore, we need to identify the defect by judging the curves of " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000569_978-3-030-20751-9_33-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000569_978-3-030-20751-9_33-Figure5-1.png", + "caption": "Fig. 5: Workspace boundary for a CDPR (Left: the projection of the desired velocity; Right: the bounding box)", + "texts": [ + " Moreover, the logistic function provides a bounded growth on the blending value to avoid unbounded large velocity commands on the CDPR. Consideration of the workspace of the CDPRs In order to guarantee a valid motion on the CDPRs, the feasible workspaces of the CDPRs have to be calculated for the teleoperation. A workspace boundary determination method [21] can compute a workspace based on different criteria, such as wrench closure or wrench feasibility, and ensure the continuity of the workspace inside a bounding box depicted in Fig. 5. Due to the continuity, the end-effector can be teleoperated in arbitrary speeds and directions based on the user\u2019s command inside the bounding box. In this method, an estimated boundary for the translational workspace with a fixed orientation is constructed by k triangular planes {F1, .., Fk}. To integrate the desired velocity with the workspace, when the endeffector reaches the triangular planes, and the direction of the desired velocity is not pointing towards the workspace (q \u2208 Fi and vT d n\u0302i > 0), the desired velocity is projected to the located plane Fi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.2-1.png", + "caption": "Figure 1.2. Rotations in R3 following various viewing angles", + "texts": [ + " Published by ISTE Ltd and John Wiley & Sons, Inc. this tool that we will perform our coordinate system transformations and position our objects in space. 1.1.1. Definition Let us recall that the j th column of the matrix of a linear application of Rn ! Rn represents the image of the jth vector ej of the standard basis (see Figure 1.1). Thus, the expression of a rotation matrix of angle ! in the plane R2 is given by: R = \" cos ! \" sin ! sin ! cos ! # . Concerning rotations in the space R3 (see Figure 1.2), it is important to specify the axis of rotation. We distinguish three main rotations: the rotation around the Ox axis, around the Oy axis and around the Oz axis. The associated matrices are respectively given by: Rx = $ % 1 0 0 0 cos !x \" sin !x 0 sin !x cos !x & ' , Ry = $ % cos !y 0 sin !y 0 1 0 \" sin !y 0 cos !y & ' , Rz = $ % cos !z \" sin !z 0 sin !z cos !z 0 0 0 1 & ' . Let us recall the formal definition of a rotation. A rotation is a linear application which is an isometry (i.e. it preserves the scalar product) and is direct (it does not change the orientation in space)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003802_safeprocess45799.2019.9213356-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003802_safeprocess45799.2019.9213356-Figure1-1.png", + "caption": "Fig. 1: Schematic diagram of an UQH", + "texts": [ + " In order to obtain the fault information, a robust three-step unscented Kalman filter (RTS-UKF), which could provide accurate fault and state estimation, is utilized. The integration of SMC and RTSUKF algorithms can guarantee that the UQH system is stable and the tracking error converges to zero asymptotically even in the presence of sensor faults. The remainder of this paper is organized as follows. The problem formulation is briefly described in Section 2. Section 3 presents the design of the proposed AFTTC scheme. Simulation results are demonstrated in Section 4. Finally, conclusions are included in Section 5. As depicted in Fig. 1, an UQH is a small multirotor helicopter, which is driven by four motors in a cross configuration. Since it has six degrees of freedom with only four control inputs, an UQH is a under-actuated system. As derived in [9], a common nonlinear dynamic model of the UQH with respect to the inertial coordinate system can be described by the following equations. x\u0308 = (cos\u03c6 sin \u03b8 cos\u03c8 + sin\u03c6 sin\u03c8)U1 m \u2212 K1 m x\u0307 y\u0308 = (cos\u03c6 sin \u03b8 sin\u03c8 \u2212 sin\u03c6 cos\u03c8)U1 m \u2212 K2 m y\u0307 z\u0308 = (cos\u03c6 cos \u03b8)U1 m \u2212 K3 m z\u0307 \u2212 g \u03c6\u0308 = Iy\u2212Iz Ix \u03b8\u0307\u03c8\u0307 + Ir Ix \u03b8\u0307\u2126r \u2212 K4 Ix \u03c6\u03072 + U2 Ix \u03b8\u0308 = Iz\u2212Ix Iy \u03c6\u0307\u03c8\u0307 \u2212 Ir Iy \u03c6\u0307\u2126r \u2212 K5 Iy \u03b8\u03072 + U3 Iy \u03c8\u0308 = Ix\u2212Iy Iz \u03c6\u0307\u03b8\u0307 \u2212 K6 Iz \u03c8\u03072 + U4 Iz (1) where x, y and z are the coordinates of the UQH with respect to the inertial frame, \u03c6, \u03b8 and \u03c8 are the roll, pitch and yaw angles, respectively, m denotes the mass of the UQH, g denotes the acceleration of gravity, Ki, i = 1, 2, \u00b7 \u00b7 \u00b7 , 6 refer to drag coefficients, Ix, Iy and Iz refer to the moment of inertia along each axis of the body-fixed frame, respectively, Ir refers to the rotational moment of inertia of the rotor, \u2126r = \u2212\u21261 + \u21262 \u2212 \u21263 + \u21264 represents the overall angular velocity of the four rotors, Ui, i = 1, 2, 3, 4 represent the control inputs, and the relationship between control inputs and angular velocities of the rotors is defined as follows: U1 = b(\u21262 1 + \u21262 2 + \u21262 3 + \u21262 4) U2 = lb(\u2212\u21262 2 + \u21262 4) U3 = lb(\u2212\u21262 1 + \u21262 3) U4 = d(\u2212\u21262 1 + \u21262 2 \u2212 \u21262 3 + \u21262 4) (2) where b is the thrust coefficient, d is the reverse moment coefficient, l is the distance between the center of the UQH and a rotor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000501_978-3-030-20131-9_238-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000501_978-3-030-20131-9_238-Figure1-1.png", + "caption": "Fig. 1: PMC schematic view.", + "texts": [ + " In Section 4, the velocity scheduling algorithm is introduced, while in Sections 5 and 6, results and their discussion are included. The model that describes the relations between the various parameters and the motion variables of the PMC, in both the Cartesian and the joint domains, is needed to conduct the necessary simulation work to validate the algorithm proposed in the sequel. This model was reported in a previous paper [16]; it is recalled here for the sake of completeness. The isostatic CRRHHRRC architecture of the PMC is illustrated in Fig. 1. It consists of two C-drives, two limbs made of one arm and one forearm, as well as the Peppermill, to which the moving plate is attached. The C-drives are differential mechanisms capable of generating translational and rotational motion of the cylindrical type independently, with the former along a direction parallel to the axis of the latter. The synchronized rotation of two nuts, driven by two parallel screws, one right-, one lefthand, attached to a DC motor, makes the controlled cylindrical motion possible", + " The MP pose is defined by the fourdimensional vector x \u2261 [ xC, yC, zC, \u03c6 ]T . Variables \u03b1i and \u03b2i, appearing in d, eq.(1b), are to be computed with the expressions: \u03b11 = arctan (h1 yC ) , \u03b12 = arctan (h2 xC ) , \u03b2i = arccos ( r2 \u2212 l2 + k2 i 2rki ) (1c) k1 = \u221a y2 C +h2 1, k2 = \u221a x2 C +h2 2, hi = zC +(\u22121)i q\u03c6 2\u03c0 , i = 1,2 (1d) The pitch of the right- and the left-hand screws of the C-drives are, respectively, p and \u2212p, while the pitch of the right- and the left-hand screws of the Peppermill are, respectively, q and \u2212q. The other geometric parameters are indicated in Fig. 1. The mathematical model used to compute the motor torques \u03c4 was obtained with the method of the natural orthogonal complement (NOC) [18]. The PMC is made up of 11 rigid bodies: four screws, two arms, two forearms, two nuts, and the Peppermill. The torque vector is computed as \u03c4 = \u03c4 d + \u03c4 s (2a) where \u03c4 d \u2261 I\u03c8\u0308 +C\u03c8\u0307 , \u03c4 s \u2261\u2212\u03b3 \u2212\u03b7 (2b) with \u03c4 \u2261 [ \u03c41L \u03c41R \u03c42L \u03c42R ]T , I \u2261 11 \u2211 i=1 Ti T MiTi, C \u2261 11 \u2211 i=1 (Ti T MiT\u0307i +Ti T WiMiTi) (2c) \u03b3 +\u03b7 \u2261 11 \u2211 i=1 Ti T (wG i +wE i ), Mi \u2261 [ Ii O O mi1 ] , Wi \u2261 [ \u03a9 i O O O ] , Ti \u2261 \u2202 ti \u2202\u03c8\u0307 (2d) where 1 and O denote the 3\u00d73 identity and zero matrices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002495_s11012-020-01147-9-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002495_s11012-020-01147-9-Figure1-1.png", + "caption": "Fig. 1 Impact geometry and contact force", + "texts": [ + "it Here, we use Hertz\u2019s theory of impact [12] along the lines already developed in [10, 11] and improve the model presented there by describing more accurately the deformation region. In addition, we investigate the effect of the impact on themodal coefficients of the beam, an aspect not addressed in [10, 11]. The paper is organized as follows: in Sect. 2 we introduce the impact model; Sect. 3 contains the numerical results and in Sect. 4 we draw our conclusions. We study the impact of a ball of mass m and radius R on a flexible beam of massM, cross section area A and length L. A fixed reference frame O\u00f0x\u0302; y\u0302\u00de is introduced on the plane of motion (see Fig. 1). Let w\u0302\u00f0x\u0302; t\u0302\u00de be the deformed beam profile and y\u0302 \u00bc z\u0302\u00f0t\u0302\u00de the vertical component of the trajectory of the ball\u2019s center of mass C, with x\u0302 2 \u00bd0; L the space variable, t\u0302 0 the time. Let \u00f0x\u0302k; z\u0302k\u00de be the coordinates of the point of impact, which we suppose to occur at t\u0302 \u00bc 0 for simplicity and let h be the slope of the beam profile at the impact point, tan h \u00bc ow ox \u00f0xk; 0\u00de: \u00f01\u00de The equation which governs the beam profile w\u0302\u00f0x\u0302; t\u0302\u00de for a homogeneous and isotropic beam of length L subject to gravity and to hinged-hinged boundary conditions is given by qA o2w\u0302 ot\u0302 2 \u00fe E J o4w\u0302 ox\u03024 \u00bc qAg \u00f02\u00de w\u0302\u00f00\u00de \u00bc w\u030200\u00f00\u00de \u00bc w\u0302\u00f0L\u00de \u00bc w\u030200\u00f0L\u00de \u00bc 0; \u00f03\u00de where q is the mass volume density of the beam, E the elastic modulus and J the area moment of inertia of the cross-section. We represent the impact by a contact force which acts in the direction normal to the beam profile; we indicate by F \u00bc F bn the force exerted by the beam on the ball and F \u00bc F bn is the force exerted by the ball on the beam, with bn the unit vector normal to the beam profile and directed upwards (see Fig. 1). According to the small displacements theory which is behind Eq. (2), we also assume that each element of the beam moves vertically only. During the impact, the spherical ball and the beam are subject to a deformation on a region D (called contact region) which we assume of circular shape with radius a\u0302, centered at the point of impact \u00f0x\u0302k; z\u0302k\u00de and lying on the upper planar face of the beam, while the contact force becomes in fact a distributed force p on the regionD. We indicate by d\u0302 the distance between the contact point of the undeformed ball and the undeformed beam profile (notional contact depth); this distance rises from zero during the compression phase of the impact, reaches a maximum value and then decreases again to zero during the restitution phase. The geometry of the impact is represented in Fig. 1 with a vertical cross-section. Both a\u0302 and d\u0302 are functions of time during the impact. From the Hertz theory of impact [12], the following relations hold among these parameters: a\u0302 \u00bc ffiffiffiffiffiffi R d\u0302 p \u00f04\u00de F \u00bc 4 3 E ffiffiffiffiffiffiffiffi R d\u0302 3 q \u00f05\u00de a\u03023 \u00bc 3F R 4E \u00f06\u00de where E is a compound elastic modulus and is given by 1 E \u00bc 1 m21 E1 \u00fe 1 m22 E2 \u00f07\u00de with m1 and m2 the Poisson\u2019s ratios and E1 and E2 the elastic moduli of the two bodies. During the impact, the contact force must be added to gravity on the r" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure8-1.png", + "caption": "Figure 8. The results of model with cross-sharp side gating system. The defects (b)(c) will not affect the casting but the only disadvantage is unstable solidification(a).", + "texts": [ + " The reason of defect is suspected to be the high speed of filling, which leading to a phenomena that not all the gas escape through the riser on time. ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 The results of linear-sharp side gating system is shown in Figure 7, which indicate that the quality of casting is satisfying and all the defects are inside the gating system that will not affect the impeller. The same thing also happens in cross-sharp side gating system. As the Figure 8. All the filling and solidification are stable, and effect-free defects in gating system and shrinkage defects in risers, the advance of side gating system is obvious. However, the solidification in cross-sharp gating system is more stable. As well as defect range in cross-sharp is smaller than the other. So it is worth for further improvements of cross-sharp gating system. The last solidification in cross-sharp gating system is zone in yellow like Figure 8(a). but this area is far away from the riser. This state indicate that the managements of risers can be performed. The Figure 9 describes the solidifications sequence of mould with more risers. The white zone means the part that solidified lastly. Only in the Figure 9(d). all the white area located in the flow ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 channel, which means the well sequence of solidification. Therefore, the best solution of gating system is cross-sharp side gating system with four risers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000997_tee.22990-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000997_tee.22990-Figure1-1.png", + "caption": "Fig. 1. Calculation model", + "texts": [ + " According to the IEEE standard, the degree of voltage unbalance is defined by the phase voltage unbalance rate (PVUR), and it can be expressed as [11]: PVUR(%) = Max [|Va \u2212 Vavg| \u00b7 |Vb \u2212 Vavg| \u00b7 |Vc \u2212 Vavg|] Vavg \u00d7 100% (1) where V a , V b and V c represent the voltage amplitude of the phases A, B and C; and Vavg = Va +Vb+Vc 3 . 2.2. Model building The basic parameters of the motor are shown in Table I. Based on the basic physical parameters of the prototype, the two-dimensional (2D) finite element analysis model is established as shown in Fig. 1. In the process of the 2D transient electromagnetic field calculation, the following assumptions are made to simplify the calculation. 1. Neglecting the displacement current, the electromagnetic field in the motor belongs to quasistationary field. 2. Materials are isotropic. 3. The permeability and conductivity of the materials are not affected by temperature. They are constant except for the stator core and the rotor core (due to magnetic saturation). 4. The end effect is represented by the inductance value in the circuit equation of the stator winding" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002043_iros40897.2019.8968006-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002043_iros40897.2019.8968006-Figure1-1.png", + "caption": "Fig. 1. Model of the acetabular cup including anchorage cone and four stimulation electrodes.", + "texts": [ + " This method is nonintrusive, which means that no change to the implementation of the deterministic model is required, and was already successfully applied in previous studies regarding bioelectrical applications [11], [12]. A preliminary version of this study has been reported in [13]. The electrostimulative hip revision system contains an acetabular and a femoral component. The acetabular component has stimulation electrodes attached to its surface, which are connected via a coil inside an insulation layer (see Fig. 1). The hemispherical stimulation electrodes have a diameter of 4 mm, while the ovoid cup has a length of 70 mm and a width of 58 mm. The acetabular cup also includes an anchorage cone, which is placed in one of five default drill holes within the cup. The femoral component used in this study is a preliminary prototype still in its design state. So far, the component is composed out of a hip stem (Hipstar, size 2)1 with a notch on one side of the implant, in which the stimulation electrode (width: 1", + "5 mm normal to the surface, and a range of 13 and 47 mm with a step size of 0.5 and 1 mm in the x- and y-component of each layer, respectively, resulting in a total of 7776 sampling points. Due to the different sizes of the areas of interest, the optimization of the beneficial stimulation volumes resulted in different stimulation amplitudes at the electrodes in both models. For the acetabular cup, the determined stimulation amplitudes are 1.704, 1.699, 1.451, and 1.534 V at electrode 1 to 4 (see Fig. 1) and 0.17 V for the femoral component. The optimization of the electrode arrangement, geometry, and stimulation amplitudes was carried out for minimum conductivity of cancellous bone, while the conductivity of bone substitute and bone marrow was set to its mean and maximum value, respectively. The uncertainty quantification of the stimulation regions was carried out for the optimized hip revision system. The optimal electrode arrangement for the acetabular cup used in this study is based on preliminary work [13]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003742_icarm49381.2020.9195386-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003742_icarm49381.2020.9195386-Figure2-1.png", + "caption": "Fig. 2. Mechanism of user adaptation.", + "texts": [ + " During walking, the assistant torque is provided by the power source. Specifically, the motor power is transmitted, in sequence, through bevel gears, Link 1, Link 2, Link 3 and finally to the front foot. The orthosis mainly has two functions. On one aspect, the device can transmit assistant torque to the ankle complex so as to aid progression motion. On the other aspect, resistant torque can be generated from the power transmission chain; it helps to rectify abnormal gait styles such as foot slap and foot drag. Fig. 2 illustrates the kinematic diagram of the device coupled with the ankle and foot complex. Link 9 and 10 are talus and front foot respectively. The adaptation of the device to users with different orientations of joint axes is mainly owing to the layout of the spherical joints. Specifically, as shown in Fig. 2, the initial axial orientations of the ankle joints of the user and the device are different. With the help of the spherical joints, mainly joints S4 and S7, the axial orientation of the device is able to change gradually and finally aligns with that of the user; the dashed and solid figures are the initial and final positions of the device respectively. Therefore, the device has the potential to adapt to users with different orientations of joint axes. III. VELOCITY TRANSMISSION RATIO In order to make the description of this section clearer, the kinematic diagram of the AFO coupled with the ankle and foot complex with only one position is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure40.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure40.2-1.png", + "caption": "Fig. 40.2 Mold box dimensions for the present investigations", + "texts": [ + " The design of part consists of various elements such as pouring basin, sprue, sprue well, and ingates, and they act as a channel to molten metal to flow into the mold cavity to get the desired shape of the mold cavity. Numerical simulation describes the applications of AutoCAST software for method design and optimization of casting process to enable to get defect-free casting. In the present investigation, simulation process of the selected CAD part was modeled in (.stl file), imported into AutoCAST for simulation. The mold box dimensions have been considered as 200 100 80 mm as shown in Fig. 40.2. The simulation was carried out by choosing the green sand casting process with aluminum-based alloy. The mold is subdivided into cubic elements with the element size of 1 mm for internal computation for uniform solidification and mold filling. The pouring temperature is considered 700 \u00b0C in casting simulation. It was observed from Fig. 40.3 that the Al\u2013Cu alloys were much prone at TC-1 region and TC-2 region. Table 40.1 Design parameters S. N. Design parameters Value (mm) Cross section 1 Pouring basin, height, length, and width 8, 20, 15 Rectangular 2 Diameter of sprue at top, bottom, and sprue length 10, 8, 23" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000755_b978-0-08-102211-5.00004-8-Figure4.26-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000755_b978-0-08-102211-5.00004-8-Figure4.26-1.png", + "caption": "Figure 4.26 Schematic illustration of the manufacturing process of APY [26]. APY, Auxetic plied yarn.", + "texts": [ + " However, the effective diameter of the APY can still be larger than its original state. Therefore the APY still exhibits auxetic behaviour. After the APY is released from the tension, both the soft yarns and stiff yarns will return to their original state due to the elasticity of the yarns. Four types of 4-ply auxetic yarns were manufactured by Ge et al. [26] by using two types of stiff yarn (polyester and polyamide filaments) and two types 120 Auxetic Textiles of soft yarn (two spandex yarns with different diameters). The manufacturing process of APYs is shown in Fig. 4.26. It includes three steps. In the first step, the two stiff yarn bobbins and the two soft yarn bobbins are alternately placed and fixed on a rotating disc. In the second step, all four yarns coming from the bobbins are twisted together due to the rotation of the disc to form the APY at point A. In the last step, the APY is taken away from the twisting area and wound onto a bobbin. In this manufacturing process, the twist, which is the most critical parameter for quality control, is adjusted by changing the rotation speed of the rotating disc and the take-up speed of the APY" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001754_icasert.2019.8934515-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001754_icasert.2019.8934515-Figure4-1.png", + "caption": "Fig. 4: Video Data Transmission Path.", + "texts": [], + "surrounding_texts": [ + "The microcontroller (Atmel ATMega328P) was programmed in the Arduino language (a subset of C++) for ease of programming. The Raspberry Pi [17] was programmed in Python as it is the official programming language for the device. The Raspberry Pi [17] works as a server and the controlling and viewing devices act as clients. Server-client communication is implemented using Transmission Control Protocol (TCP) for secure and reliable data transmission. We used two separate webservers for bot/arm control and live video streaming. The server for control was programmed in Python and the live video streaming server was set up using LIGHTTPD [24], which is a light, flexible, fast and less resource-consuming webserver. Single-board computers like the Raspberry Pi [17] are usually low on hardware resources and as such, we need lightweight software to compensate for the hardware limitations. Additionally, CRTMPSERVER [25] was set up for temporarily storing the video data coming from the camera module [21]. In order to display the live video, a webpage was required. PHP and HTML languages were used to create the webpage, which is hosted on the LIGHTTPD [24] webserver. We used STROBE Media Playback software on the webpage to play the live video. The bot software, i.e. the control app establishes a central control system to control the bot from a remote distance with the help of an android device. It offers a user-friendly Graphical User Interface (GUI), which includes all the necessary tools and information display areas." + ] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure1.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure1.2-1.png", + "caption": "Fig. 1.2 STL file faceted approximation of CAD geometry (Source Oak Ridge National Lab.)", + "texts": [ + " process is automated, variation introduced from operator interactions during the pre-processing and post-processing operations have significant impacts on the final component characteristics. The acceptable amount of variability depends on how critical the performance expectations of the component are. While the individual processes vary significantly in their materials and processing methodology, the framework and software is universal. The industry accepted file format for AM is the STL, short for Stereolithography which was developed in the early 1980s. This format represents a computer aided drafting (CAD) model\u2019s geometry by faceted surfaces as shown in Fig. 1.2. This geometric model is virtually \u201csliced\u201d into layers and used to generate deposition paths for each layer of the component. Each layer is deposited sequentially on top of the previous layers to form the finished component. The production process flow is shown pictorially in Fig. 1.3. Support material is removed from locations with overhangs and finishing operations are performed tomeet the specifications on geometry, surface quality and/or resolution. Often these finishing operations involve sanding, vapor distillation smoothing, or machining" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000754_ilt-09-2018-0356-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000754_ilt-09-2018-0356-Figure1-1.png", + "caption": "Figure 1 Mechanical layout of standard FZG back-to-back gear test rig (left) and FZG efficiency gear test rig (right)", + "texts": [ + " The experimental investigations were conducted on FZGbackto-back gear test rigs under dip-lubrication. The scuffing load capacity and pitting lifetime were investigated on the FZG back-to-back gear test rig according to DIN ISO 14635- 1:2006-05 (2006), whereas the friction behavior was investigated on a modified FZG back-to-back gear test rig \u2013 the FZG efficiency gear test rig. The following descriptions of the main features of the test rigs are based on the works and formulations of Lohner et al. (2015a) and Ziegltrum et al. (2017). Figure 1 shows the mechanical layout of the standard FZG back-to-back gear test rig and the modified FZG efficiency gear test rig. Both test rigs are based on the principle of power circulation, whereby gear pairs with identical gear ratio are mounted in test and slave gear box. A load clutch applies the transmitted torque. The electric engine only has to supply the losses of test and slave gear box. To provoke gear failures in the test gear box, helical gears with a higher load carrying capacity aremounted in the slave gear box of the standard FZG standard back-to-back gear test rig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002922_icieam48468.2020.9112020-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002922_icieam48468.2020.9112020-Figure7-1.png", + "caption": "Fig. 7. , b \u2013 Initial positions of the stages: N=11, b N=12, c \u2013 Final position of the stages N=12.", + "texts": [ + " Downloaded on June 26,2020 at 08:14:26 UTC from IEEE Xplore. Restrictions apply. step ( 10 2 k O P ).The vector of generalized coordinates at this stage is written as follows 10 1N q , and the equations of constraints have the form: 10 10 1 1 mincosN Nl l , (36) 10 10 10 * 2 2 3 maxcosN N Nl l l l , (37) 10 10 10 10 1 1 2 2sin sinN N N N Pl l h The next stage of motion N=11 occurs at the supports 3 and 4, fixed on the surface of the step ( 11 3,4 constN O r , 11 3 min Nl l , 11 3 0N ), as shown in Fig. 7, . Upon this, link 2 is parallel to the plane , its length is minimal ( 11 2 constN O r , 11 2 min Nl l , 11 2 0N ). At the specified stage, the rotation of link 1 occurs in the vertical plane 11 1 1( )N t relative to the fixed support 2, as well as the reduction of its length 11 1 1( )Nl l t until the link is horizontal 11 1 0k , and its length will be minimal The vector of generalized coordinates is defined as follows T 11 1 1( , )N l q The final stage of motion N=12 occurs in the horizontal plane of the stage, the initial position of the robot at this stage is shown in Fig. 7, b. At this stage, three supports are fixed on the surface 2 \u2013 4 ( 12 2 4 constN O r ), all the links have the minimum length 12 1 3 min Nl l and are located in the horizontal plane of the step 12 1 3 0N . The stage consist in rotating the link 1 in the horizontal plane 12 1 1( )N t counterclockwise (the direction of rotation of link 1 is shown in Fig.7, b) as long as the support 1 will contact the surface of the step, thus the condition for completing the stage can be written as follows: 12 1 k ! (Fig. 7, c). The vector of generalized coordinates at this stage is written as 12 1N q At this point, the algorithm of the robot\u2018s crawling onto the step of the flight of stairs is finalized. All the links of the robot are located on the surface of the step. It should be noted that the proposed algorithm can be used to overcome such a flight of stairs, the ratio of the step sizes with the lengths of the links of the robot can be described as follows: 2 2 max max minPh l l min minPh l min maxPl l V. NUMERICAL SIMULATION RESULTS To verify the adequacy of the proposed algorithm for the robot\u2019s crawling onto the step of the flight of stairs, a software package has been developed that allows for the numerical simulation of the movement of the specified device" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002106_tpel.2020.2971226-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002106_tpel.2020.2971226-Figure9-1.png", + "caption": "Fig. 9. Sizes of the commutator segments.", + "texts": [ + " All the stages in the reverse commutation process are symmetric to those in the forward commutation process, and the capacitor C is discharged to zero voltage after the reverse commutation process. Each phase alternates in the forward and reverse commutation processes so that the capacitorC is charged and discharged periodically. Experiments were carried out to verify the proposed hybrid commutation theory. A prototype of the hybrid dc motor is shown in Fig. 8, the sizes of the commutator segments are shown in Fig. 9, and the parameters are shown in Table I. In an experiment of a brushed dc motor, it is difficult to directly measure the current of the commutating coil, so the parameters, such as the contact voltages of the brushes and voltages of the commutator segments, are measured to investigate the commutation process [33], [34]. As in the case of a brushed dc motor, the windings of the hybrid dc motor rotate together with the rotor, so the phase voltages and currents cannot be directly measured. Therefore, the voltages and currents of the commutator segments were measured in these experiments, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001557_ias.2019.8912014-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001557_ias.2019.8912014-Figure6-1.png", + "caption": "Fig. 6. Discharge end of prototype shuttle car", + "texts": [ + " Therefore, the body was divided into four major parts, plus 10 additional components to be cemented/bolted together. The majority of parts were printed on a Gigabot 3+ 3D printer, which has a 600 x 600 mm print bed, wide enough to print the full width of each body part. Smaller parts were printed on a Makerbot Replicator Z18 3-D printer which has a 300x300 mm print bed. In addition to the body parts, two additional frame rails were 3-D printed to provide clearance for the body to sit above the steering servo motors. Fig. 6 shows the discharge end of the prototype and Fig. 7 shows the load end of the prototype mounted on the frame. Fig. 8 shows the completed model. Fig. 7. Load end of prototype shuttle car Page 4 of 7 978-1-5386-4539-0/19/$31.00 \u00a9 2019 IEEE 2019-MIC-0825 Table II shows the laboratory prototype dimensions compared with the full-size shuttle car. This table shows that a scale factor of approximately 1:6 was preserved for all dimensions except for the tires. Because of the uniqueness of the axles selected for the prototype, no wheel/tire combinations could be found at a 1:6 scale" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001786_s11837-019-03963-1-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001786_s11837-019-03963-1-Figure2-1.png", + "caption": "Fig. 2. The complex unit circle in the f-plane maps onto hypotrochoidal holes in the Z-plane.", + "texts": [ + " The parameter c controls the sharpness of the points and can take values between 1 3 and 1 3, inclusive. Special values include c \u00bc 0, which defines the unit circle in the Z-plane (an identity mapping), and c \u00bc 1 3, which leads to cusped points. Values of c outside this range lead to self-intersection. Note that negative values of c yield the same shapes as positive values, but with the shape rotated by 45 . The void shape can also be written in parametric form as x \u00bc cos/\u00fe c cos 3/ y \u00bc sin/ c sin 3/: \u00f02\u00de Several hypotrochoids with various values of c are plotted in Fig. 2. To find the in-plane moduli, the stress and displacement fields around a single void must first be found using the complex-variable method of elasticity. Two loading cases are considered: 1. Uniaxial loading 2. Shear loading The uniaxial loading case is used to find the effective inplane Young\u2019s modulus Exx, while the shear loading case is used to find the effective in-plane shear modulus Gxy. Both loading conditions are applied far away from the void (i.e., at infinity). The void is surrounded by an infinitely large matrix material with isotropic properties E and m" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001229_systol.2019.8864774-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001229_systol.2019.8864774-Figure1-1.png", + "caption": "Fig. 1. Reference frames of hexarotor.", + "texts": [ + " The paper is organized as follows: the mathematical model of the multirotor UAV, the propulsion system and the compu- 978-1-7281-0380-8/19/$31.00 \u00a92019 IEEE 104 tation of energy consumption are presented in Section II. The main elements defining the computation of flight endurance and maximum energy are described in Section III. The effect of fault in actuators is presented in Section IV. Finally, the simulation results as well as conclusions and future works are presented in Section V and VI. The hexarotor is composed by six BLDCMs attached to a rigid cross frame as shown in Fig. 1. The movement of the hexarotor is performed around two reference frames: the fixed inertial frame, denoted by Ee, and the non-inertial frame or rigid body frame, denoted by Eb. [x, y, z]T is the position vector of the center of mass of the hexarotor relative to the fixed inertial frame Ee and [\u03c6, \u03b8, \u03c8]T is the orientation vector described by the Euler angles: \u03c6 (roll) angle around xaxis, \u03b8 (pitch) around y-axis, and \u03c8 (yaw) around z-axis. The dynamical model is given by the equations set (1) according to [19] as, x\u0308 = (c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8) U1 m y\u0308 = (c\u03c6s\u03b8s\u03c8 \u2212 s\u03c6c\u03c8) U1 m z\u0308 = \u2212g + (c\u03c6c\u03b8) U1 m \u03c6\u0308 = Iy \u2212 Iz Ix \u03b8\u0307\u03c8\u0307 + Jm Ix \u03c9T \u03b8\u0307 + U2 Ix \u03b8\u0308 = Iz \u2212 Ix Iy \u03c6\u0307\u03c8\u0307 \u2212 Jm Iy \u03c9T \u03c6\u0307+ U3 Iy \u03c8\u0308 = Ix \u2212 Iy Iz \u03b8\u0307\u03c6\u0307+ U4 Iz , (1) where s and c are the trigonometric functions sin and cos, m is the total mass, g is the acceleration due to the gravity, Jm is the total inertia of the motors, Ix, Iy and Iz are the inertial moments around the x, y, z axis, respectively", + " (13) From (13), the angular speed of each BLDCM can be computed as: \u03c9i = \u221a T bNm , (14) by considering the mechanical model of BLDCM (5) and the relationship between current and duty cycle (6), the total current demanded by all BLDCM is computed as: Ii = ( d\u03c92 i \u2212Df\u03c9i \u2212 Tfric ) Dci KE . (15) Finally, considering the existing relationship between the battery C-rate (11) and the total current, the flight endurance is computed using equation (12). Note that equation (15) represents the total current in hover condition and it allows to compute the maximum flight endurance according to magnitude of battery capacity. From Fig. 1 the actuators (BLDCMs) are distributed on hexarotor frame in a \u201dX\u201d configuration where the front of the vehicle is goings to x-axis. According to [25], the control allocation and re-allocation method can be used to accommodate actuator faults. In addition, the fault tolerant control can be achieved without the need to reconfigure/restructure the baseline controller due to the control signals are redistributed among faulty and healthy actuators [26]. Control allocation method is usually applied on over-actuated systems where the amount of actuators is greater than the controlled variables, as it can noted in relationship between the square of angular speed of each BLDCM and control inputs actuating on hexarotor dynamics defined in (2)", + " The hexarotor develops a mapping mission which comprises to flight over an area of 8000 m2 inside of a safety region given a total region area of 12000 m2. The total trajectory is obtained considering the approach given in [28]. In Fig. 3 the result of the tracking controller with the reference path in 3D view is shown. In the nominal case, the total energy consumption is 172.5 kJ or 47.92 Wh which is in the range of the maximum energy that battery can supply, and a maximum flight time of 14.08 min. The fault effect was simulated as a loss effectiveness of 70 % of BLDCM 1 at 200 sec and 30 % of BLDCM 5 at 400 sec (according to Fig. 1). In Fig. 4 and Fig. 5 the result of the battery current and battery voltage for nominal and faulty cases is shown. When the fault occurs the battery current increase due to the tracking controller try to compensate the fault effect and it can be also observed that the discharge rate increase. This last can be noted in the State of Charge in Fig. 6 where at the end of the mission the minimum SoC in nominal case is 50.24 % and in faulty case is 39.48 % with a final energy of 207.3 kJ or 57.58 Wh and maximum flight endurance of 18 min and 13 sec" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure82.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure82.5-1.png", + "caption": "Fig. 82.5 SS GCI and AlSiC profile of Von Mises stress (Model 2)", + "texts": [], + "surrounding_texts": [ + "In the coupled analysis, the thermal load was coupled with the structural load to find out the combined effect on brake disc models. The temperature induced at various time points was imported into static structural, and then structural loads were applied. The analysis was run for 36 s, i.e. the time taken by the vehicle to stop due to the application of the emergency brake. The output results of von Mises stress and total deformation developed in the model were recorded (Figs. 82.7, 82.8 and 82.9). Some important points that can be drawn from the analysis are: 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 979 \u2022 When compared to the Factor of Safety offered by the GCI models, the AlSiC models offer higher Factor of Safety. \u2022 For the same applied load, the AlSiC models have lower thermal stresses than the GCI models, as AlSiC material has greater thermal conductivity and heat dissipation capability. \u2022 The weight of AlSiC models is lesser when compared to the GCI models (GCI model having a weight of about 134 kg gets reduced to 54 kg in case of AlSiC material). 980 E. Madhusudhan Raju et al." + ] + }, + { + "image_filename": "designv11_80_0001959_s0036029519130111-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001959_s0036029519130111-Figure2-1.png", + "caption": "Fig. 2. Stress intensity fields. Contact with the substrate.", + "texts": [ + "4 according to Siebel. The temperature problem was not solved in parallel. The substrate made an angle of 2\u00b0 with the horizontal plane, which imitates the real tool inclination during coating deposition, which is used in practice. When the tool is in contact with the substrate, the intensity of internal stresses increases to the ultimate tensile strength of the material at room temperature (300\u2013310 MPa) at the lower edge under the action of an axial load and friction at the tool\u2013substrate boundary (Fig. 2). As a result, elements fail according to the fracture criterion used in the calculation, and the total number of elements in the mesh decreases. The critical stress zone is small as compared to the total tool volume. Therefore, the material fails on the friction surface in terms of the criteria and assumptions used in the model. The calculation results demonstrate that deviation of tool from the given axis of rotation is small, 0.6 mm per radius, at a tool length of 40 and 60 mm and a diameter of 20 mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001596_ecce.2019.8912720-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001596_ecce.2019.8912720-Figure1-1.png", + "caption": "Fig. 1. Sectional view of SynRM: (a) d-axis flux path, (b) q-axis flux path.", + "texts": [ + " On the other hand, exponential functions were used in modeling the self and cross saturation effects [10]. However, both approximation methods did not offer an explicit MTPA solution. In [11], the flux was approximated by Taylor series, and the MTPA points were obtained as the solution of a fourth order polynomial. In this work, the flux maps are modeled analytically by the 2-D Taylor series approximation. The coefficients are obtained by the least square method. With the computed cross coupled inductances, the MTPA points are calculated analytically. II. FLUX MODEL AND TORQUE EQUATION Fig. 1 shows SynRM section diagram with flux lines. The rotor has multi-layer flux barriers so that magnet saliency is polarized depending on the direction. However, the SynRM flux can not be described by a linear model due to the core saturation. Fig. 2 shows constant torque lines based on a linear model (bold solid lines). Note that they are symmetric with respect to the 45\u25e6 line. However, the actual constant torque curves appear asymmetrically (thin solid lines). Note that the MTPA points are determined at the intersection between a constant torque curve and a circle" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003925_ecce44975.2020.9235775-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003925_ecce44975.2020.9235775-Figure9-1.png", + "caption": "Fig. 9. Machine geometry and winding layout for study machine.", + "texts": [ + " In this paper, an accurate dynamic model for a machine with ISC is implemented in MATLAB/Simulink. In [16], the authors of this paper presented a precise model for torque and failure path current waveforms prediction. In this paper, such a model is combined with the voltage and flux linkage equations. Thus, a full model of the machine can be analyzed. A block diagram of the proposed model in this paper is shown in Fig. 8. To validate this model, FEA results are compared with the proposed model\u2019s results. The geometry implemented in FEA is shown in Fig. 9, with rated values and parameters described in Table II. Each phase contains six coils, and each coil has 33 turns. The machine to analyze has been designed with the same stator cross-section of the machine tested in the previous sections. Therefore, it is assumed that the thermal limit of the stator is reached at the stall current of 4.8 pu. Without loss of generalization, 15 turns were shorted in coil 1 operating at rated conditions with current control. The FEA results are compared with the proposed model in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001748_s1052618819060074-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001748_s1052618819060074-Figure2-1.png", + "caption": "Fig. 2. Mechanism of the Delta family with four degrees of freedom.", + "texts": [ + " The main advantage over the classic schemes is the higher speed of printing without losing quality. Despite the fact that these two variants of the Delta-scheme have been studied well [5\u20138] and various ways for improving their technological parameters have been proposed, the modifications of the mechanism with different numbers of degrees of freedom have been studied only a little. One of the possible modifications of Delta-robot is a mechanism for which the scheme and the photo of a real prototype of which is shown in Fig. 2. Using four linear drives in conjunction with the separation of one of the chains with the parallelogram into two chains with a PUS-structure (P, prismatic pair; U, Hooke joint; S, spherical hinge) or PUU-structure makes it possible to implement an additional rotational degree of freedom (around the axis, parallel to axis y, and passing through the output link). In more detail, the features of the structure and kinematics of the mechanism under consideration are described in [9]. Under parametric synthesis of any parallel mechanism, it is necessary to investigate the possibility of special positions arising [10\u201313], as when the mechanisms ingress into these positions, the mobility of 503 output link is changed (the loss of a degree of freedom or uncontrollability arise) [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002215_s00170-019-04806-8-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002215_s00170-019-04806-8-Figure5-1.png", + "caption": "Fig. 5 Projections of velocities of the contact points during the rotation around hole\u2019s axis onto axes \u041e1\u0445 and \u041e1\u0443", + "texts": [ + "3) \u03b51\u2014angle between segment \u0412L and axis of hole, \u03b52\u2014angle between segment \u0412L and peg axis, cos\u03b51 \u00bc a1 OB; sin\u03b51 \u00bc S1 OB; sin\u03b52 \u00bc S2 OB; cos\u03b52 \u00bc a2 OB; B \u00bc OB \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a21 \u00fe S21 q \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi a22 \u00fe S22 q ; BL = 2OB = 2B. Projections of velocities V\u03b3 B1; V\u03b3 B2; V \u03b3 K to fixed axes\u041e1\u0445, \u041e1\u0443, and \u041e1z (Fig. 4) are equal to V\u03b3 B1x \u00bc V\u03b3 B2x \u00bc \u2212V\u03b3 B1hcos\u03c8 \u00bc \u22122a1cos\u03c8\u03b3 ; V\u03b3 K\u0445 \u00bc \u2212KLsin\u03b3cos\u03c8\u03b3 ; V\u03b3 B1y \u00bc V\u03b3 B2y \u00bc \u2212V\u03b3 B1hsin\u03c8 \u00bc \u22122a1sin\u03c8\u03b3 ; V\u03b3 Ky \u00bc \u2212KLsin\u03b3sin\u03c8\u03b3 ; V\u03b3 B1z \u00bc V\u03b3 B2z \u00bc \u22122S1 \u03b3 : V\u03b3 Kz \u00bc KL \u03b3 cos\u03b3: \u00f012\u00de Rotation about hole axis occurs with angular velocity of \u03c8\u0307 \u00bc d\u03c8 dt . Velocities of contact points in this movement are located in the fixed plane \u041e1\u0445\u0443, directed (see Fig.5) at tangents to the hole aperture edge circumference and at all contact points are equal in magnitude V\u03c8 B1 \u00bc V\u03c8 B2 \u00bc V\u03c8 K \u00bc 0:5D\u03c8\u02d9 : Projections of these velocities to the fixed axes of coordinates \u041e1\u0445 and \u041e1\u0443 (Fig. 5) taking values of sin\u03c4 \u00bc b 0:5D and cos\u03c4 \u00bc S1 0:5D into account shall be transformed to the following form: V\u03c8 B1x \u00bc \u2212 S1sin\u03c8\u2212bcos\u03c8\u00f0 \u00de\u03c8 ; V\u03c8 B2x \u00bc \u2212 S1sin\u03c8\u00fe bcos\u03c8\u00f0 \u00de\u03c8 ; V\u03c8 Kx \u00bc 0:5Dsin\u03c8\u03c8 ;V\u03c8 B1y \u00bc S1cos\u03c8\u00fe bsin\u03c8\u00f0 \u00de\u03c8 ; V\u03c8 B2y \u00bc S1cos\u03c8\u2212bsin\u03c8\u00f0 \u00de\u03c8 ;V\u03c8 Ky \u00bc \u22120:5Dcos\u03c8\u03c8 ; V\u03c8 B1z \u00bc 0;V\u03c8 B2z \u00bc 0;V\u03c8 Kz \u00bc 0: \u00f013\u00de To define projections of velocities V\u03c8 B1 and V \u03c8 B2 onto moving axes of coordinates\u041e2\u03b5 and\u041e2\u03b7, they are first divided into two components. One of these components is parallel to axis \u041e1\u0435, and the second one is parallel to axis \u041e1h (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000804_j.trpro.2019.07.074-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000804_j.trpro.2019.07.074-Figure3-1.png", + "caption": "Fig. 3 Cutting section of 3D model geometry", + "texts": [], + "surrounding_texts": [ + "The aim of the analysis was to examine and analyze the oscillation of the investigated object, which was rolling ball bearing (see Fig. 2, 3) connected to the rotary shaft in the test center (see Fig. 4). The measurement itself was carried out in several steps to examine and compare the individual bearings with respective deformations in the orbits. The basis of the measurement was to obtain data on the dynamic behavior of individual bearings without and with deformations in orbits for the need to detect vibrations and to understand the behavior of the examined mechanical components during operation. Measurements were made at the start-up of the test device (with appropriate bearing on the rotary shaft) at 1000 rpm, staying at this level of mentioned rpm and runout of testing machinery in a total cycle of approximately 50 s. 514 Peter Sulka et al. / Transportation Research Procedia 40 (2019) 511\u2013518 4 Sulka et al./ Transportation Research Procedia 00 (2019) 000\u2013000 The measurement conditions were as follows: a. assembly of sensors and measuring instruments (see Fig. 5), b. performing measurements at a defined time (see Fig. 5), c. do not exceed the maximum speed of 1000 rpm, d. compliance with security measures." + ] + }, + { + "image_filename": "designv11_80_0001250_demped.2019.8864901-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001250_demped.2019.8864901-Figure2-1.png", + "caption": "Fig. 2. Schematic of circuit to make electrolytic corrosion bearing.", + "texts": [ + " Since the local stress concentration is induced by the formation of recesses, the progress of deterioration is accelerated. The electrolytic corrosion occurs when the current in the interior of the bearing during rotation has passed. It sparks through a very thin oil film on the rolling contact portion, the surface is locally dissolved, and it makes what at first glance appears to be a pitching form. This time, the electrolytic corrosion bearing is artificially made at the load side, as shown in Fig. 2. The opposite side bearing was replaced with an insulated bearing. Then, only the DC current passed through the load side bearing while operating the motor. By exchanging the opposite load side with an insulated bearing, it is possible to corrode only the load side bearing. Generally, by energizing the electric current inside the bearing during operation, a spark occurs through the thin oil film at the rolling contact part, and an uneven surface shape is formed. The electrolytic corrosion bearing was created by flowing a direct current of 10 A through load side bearing for about 60 minutes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000316_978-981-13-6469-3_14-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000316_978-981-13-6469-3_14-Figure5-1.png", + "caption": "Fig. 5 Outer race fault", + "texts": [ + " For the vibration measurement and analysis, four channels FFT analyser is used (Fig. 3). The three piezoelectric accelerometers are used to collect the vibration signals in three directions at a bearing-housing. The ball bearing geometry is shown in Fig. 4. The self-aligning, double row deep groove ball bearing (SKF 1205 ETN) is used with the dimensions shown in Table 2. The fault in the Inner race, outer race and cage created using electro-discharge machining (EDM), the rectangular notch of 7.5 mm length, width 0.55 mm and depth 1.15 mm is created in outer race (Fig. 5). Similarly, rectangular notch of 7.5 mm length, width 0.5 mm and depth 1.15 mm, created in the inner race (Fig. 6), and rectangular notch of 6 mm length, width 0.7 mm and depth 1 mm, in cage (Fig. 7). The faulty ball (spall) from other bearing was used for analysis (Fig. 8). The signal is filtered in (DEMOD fmin, DEMOD fmax) range and demodulated. Parameters for G-demod wideband measurement is as shown in Table 1. The specification of rolling ball bearing as shown in Table 2. The bearing characteristic frequency depends on the geometry of the bearing and on the type bearing defect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure1-1.png", + "caption": "Figure 1. Plain journal bearing complement textured", + "texts": [], + "surrounding_texts": [ + "velocities,\ufeffwhile\ufeffshowing\ufeffthe\ufeffdisplacement\ufeffevolution\ufeffwhich\ufeffis\ufeffdue\ufeffto\ufeffthe\ufeffhydrodynamic\ufeffpressure\ufeffeffect,\ufeff deformation\ufeffas\ufeffwell\ufeffas\ufeffshear\ufeffstresses\ufeffusing\ufeffthe\ufeffANSYS\ufeffWorkbench\ufeffcalculation\ufeffcode,\ufeffby\ufeffsolving\ufeffthe\ufeff displacement\ufeffequations\ufeffusing\ufeffthe\ufefffinite\ufeffvolume\ufeffmethod.\nThe\ufeffmechanical\ufeffdisplacements\ufeffof\ufeffthe\ufeffcontact\ufeffsurfaces\ufeffaffect\ufeffthe\ufeffoil\ufefffilm\ufeffgeometry\ufeffand\ufefftherefore\ufeff the\ufeffperformance\ufeffof\ufeffthe\ufeffhydrodynamic\ufeffplain\ufeffbearing.\ufeffThe\ufeffdisplacements\ufeffthat\ufeffwe\ufeffwill\ufeffpresent\ufeffin\ufeffthis\ufeff study\ufeffare\ufeffgeometrical\ufeffmodifications,\ufeffwhich\ufeffoccur\ufeffduring\ufeffthe\ufeffoperation\ufeffof\ufeffthe\ufeffplain\ufeffbearing,\ufeffas\ufeffwell\ufeff as\ufeffthe\ufeffshear\ufeffstresses.\ufeffThese\ufeffelastic\ufeffdisplacements\ufeffwhich\ufeffare\ufeffdue\ufeffto\ufeffthe\ufeffhydrodynamic\ufeffpressure\ufefffield\ufeff modify\ufeffthe\ufeffgeometrical\ufeffshape\ufeffof\ufeffthe\ufeffplain\ufeffbearing\ufeffby\ufeffinducing\ufeffan\ufeffincrease\ufeffor\ufeffa\ufeffreduction\ufeffof\ufeffthe\ufeffoil\ufeff film\ufeffthickness.\nBACKGRoUNd\nBouyer\ufeffpresented\ufeffin\ufeff2003\ufeffan\ufeffexperimental\ufeffand\ufeffnumerical\ufeffstudy\ufeffon\ufeffthe\ufeffthermal\ufeffand\ufeffelastic\ufeffeffect\ufeff on\ufeffthe\ufeffhydrodynamic\ufeffperformances\ufeffof\ufeffa\ufeffnon-textured\ufeffmisaligned\ufeffbearing.\ufeffThis\ufeffresearch\ufeffshows\ufeff that\ufeffthe\ufeffmisalignment\ufeffvery\ufeffsignificantly\ufeffmodifies\ufeffthe\ufeffoperating\ufeffparameters\ufeffof\ufeffthe\ufeffbearing.\ufeffHe\ufeff also\ufefftested\ufeffa\ufeffworn\ufeffbearing\ufeffthat\ufeffis\ufeffto\ufeffsay\ufeffa\ufeffdesired\ufeffgeometric\ufeffmodification\ufeffon\ufeffthe\ufeffperformance\ufeff of\ufeff the\ufeff bearing.\ufeff These\ufeff results\ufeff show\ufeff that\ufeff in\ufeff the\ufeff case\ufeff of\ufeff a\ufeff very\ufeff loaded\ufeff bearing,\ufeff the\ufeff generated\ufeff displacements\ufeffmay\ufeffbe\ufeffof\ufeffthe\ufefforder\ufeffof\ufeffthe\ufeffminimum\ufeffthickness\ufeffof\ufeffthe\ufeffoil\ufefffilm,\ufeffand\ufeffthe\ufeffheating\ufeffof\ufeff the\ufefflubricant\ufefffilm\ufeffleads\ufeffto\ufeffa\ufeffsignificant\ufeffdecrease\ufeffin\ufeffviscosity.\ufeffFluid\ufeffand\ufeffat\ufeffa\ufeffdifferential\ufeffexpansion\ufeff of\ufeffthe\ufeffelements\ufeffof\ufeffthe\ufeffnon-textured\ufeffbearing.\nBendaoud\ufeffet\ufeffal\ufeffshow\ufeffin\ufeff2006\ufeffthe\ufeffeffect\ufeffof\ufeffthe\ufeffpressure\ufefffield\ufeffon\ufeffthe\ufeffstresses\ufeffdistribution\ufeffin\ufeff an\ufeffuntextured\ufeffhydrodynamic\ufeffjournal\ufeffbearing.\ufeffThis\ufeffstudy\ufeffis\ufeffcarried\ufeffout\ufeffby\ufeffsolving\ufeffdisplacement\ufeff equations\ufeffusing\ufeffthe\ufefffinite\ufeffelement\ufeffmethod.\ufeffTheir\ufeffresults\ufeffshow\ufeffthat\ufeffelastic\ufeffdeformations\ufeffare\ufeff important\ufeffand\ufeffpreponderant\ufefffor\ufeffhigher\ufeffrotational\ufeffspeeds,\ufeffand\ufeffshear\ufeffstresses\ufeffare\ufeffimportant\ufeff for\ufeff higher\ufeff hydrodynamic\ufeff pressures.\ufeff In\ufeff 2014,\ufeff Bendaoud\ufeff conducted\ufeff a\ufeff search\ufeff on\ufeff behavior\ufeff elastohydrodynamic\ufeffplain\ufeffbearing\ufeffhydrodynamic\ufeffvery\ufeffloaded.\ufeffMehala\ufeffet\ufeffal\ufeff treats\ufeff in\ufeff2016\ufeff the\ufeffeffect\ufeffof\ufeffthe\ufeffnon-Newtonian\ufeffflow\ufeffPerformance\ufeffand\ufeffthermal\ufeffeffect\ufeffon\ufeffa\ufeffbearing\ufeffcoated\ufeff with\ufeffa\ufeffhigh\ufefftin\ufeffcontent.\nIn\ufeff 2016,\ufeff Shahab\ufeff Hamdavi\ufeff and\ufeff collaborator,\ufeff presents\ufeff the\ufeff effect\ufeff of\ufeff partially\ufeff textured\ufeff surface\ufeff of\ufeffhydrodynamic\ufefflong\ufeffjournal\ufeffbearing\ufeffon\ufeffthe\ufeffpressure\ufeffdistribution\ufeffand\ufeffload\ufeffcarrying\ufeffcapacity\ufeffwas\ufeff studied.\ufeffThe\ufeffequations\ufeffof\ufeffpressure\ufeffdistribution\ufeffand\ufeffload\ufeffcarrying\ufeffcapacity\ufeffwere\ufeffderived\ufeffand\ufefftheir\ufeff trends\ufeffwere\ufeffrepresented.\ufeffThe\ufeffresults\ufeffshow\ufeffthat,\ufeffapplying\ufeffpartial\ufeffsurface\ufefftexture\ufeffhad\ufeffa\ufeffpositive\ufeffand\ufeff remarkable\ufeffeffect\ufeffon\ufefffunctional\ufeffcharacteristics\ufeffof\ufeffhydrodynamic\ufeffjournal\ufeffbearings,\ufeffi.e.\ufeffload\ufeffcarrying\ufeff capacity\ufeffand\ufefflubricant\ufefffilm\ufeffpressure.\nRufei\ufeffet\ufeff al\ufeff treated\ufeff in\ufeff2016,\ufeffanalyzed\ufeff the\ufeff influences\ufeffof\ufeff flexibility\ufeffand\ufeffsurface\ufeff texture\ufeffon\ufeff the\ufeff performances\ufeffof\ufeffa\ufefffinite-long\ufeffjournal\ufeffbearing.\ufeffReynolds\ufeffequation\ufeffhas\ufeffbeen\ufeffsolved\ufeffby\ufeffusing\ufefffinite\ufeff difference\ufeffmethod\ufeffand\ufeffWinkler\ufefffoundation\ufeffmodel\ufeffhas\ufeffbeen\ufeffused\ufeffto\ufeffcalculate\ufeffthe\ufeffelastic\ufeffdeformation\ufeff of\ufeffthe\ufeffbearing\ufeffbush\ufeffor\ufeffhousing.\ufeffThey\ufeffare\ufeffindicated\ufeffthat\ufeffthe\ufeffsurface\ufefftexture\ufeffcan\ufeffappreciably\ufeffchange\ufeff the\ufeffshaft\ufeffcenter-line\ufefflocus\ufeffand\ufeffoil\ufefffilm\ufeffpressure\ufeffdistribution\ufeffcomparatively\ufeffto\ufeffthe\ufeffsmooth\ufeffcase.\ufeffThe\ufeff texture\ufeffarrangements\ufeffon\ufeffthe\ufeffbearing\ufeffsurface,\ufeffthe\ufefftexture\ufeffnumber\ufeffaccording\ufeffthe\ufeffcircumferential\ufeffaxis\ufeff or\ufeffaxial\ufeffaxis,\ufeffalso\ufeffthe\ufefftexture\ufeffdepth\ufeffin\ufeffthe\ufeffradial\ufeffdirection,\ufeffall\ufeffof\ufeffthem\ufeffhave\ufeffsignificant\ufeffinfluences\ufeffon\ufeff the\ufeffstatic\ufeffcharacteristics\ufeffof\ufeffa\ufeffflexible\ufeffjournal\ufeffbearing.\nIn\ufeff2018,\ufeffZhang\ufeffet\ufeffal.\ufeffpresented\ufeffresearch\ufeffon\ufeffthe\ufeffeffect\ufeffof\ufeffthe\ufeffthermal\ufeffand\ufeffelastic\ufefffor\ufeffmisaligned\ufeff bearing\ufeffwith\ufefftexture\ufeffsurface.\ufeffThe\ufeffbearing\ufeffis\ufeffsubjected\ufeffunder\ufeffhigh-speed\ufeffand\ufeffheavy-load\ufeffconditions.\ufeff As\ufeffwell,\ufeffthey\ufefftake\ufeffon\ufeffconsideration\ufeffthe\ufeffeffects\ufeffof\ufeffcavitation\ufeffand\ufeffviscosity-temperature.\ufeffZhang\ufeffand\ufeff collaborators\ufeff show\ufeff that\ufeff the\ufeff maximum\ufeff oil\ufeff film\ufeff pressure\ufeff and\ufeff load-carrying\ufeff capacity\ufeff of\ufeff a\ufeff journal\ufeff bearing\ufeffincrease\ufeffwith\ufeffjournal\ufeffmisalignment.\ufeffMeanwhile,\ufeffthe\ufeffoil\ufefffilm\ufefftemperature\ufeffincreases\ufeffsharply\ufeff while\ufeffconsidering\ufeffmisalignment.", + "THEoRETICAL ANALySIS\nThe\ufeffpressure\ufefffield\ufeffis\ufeffdetermined\ufeffby\ufeffthe\ufeffresolution\ufeffof\ufeffthe\ufeffgeneralized\ufeffNavier-Stokes\ufeffequation\ufeffunder\ufeff classic\ufeffassumptions\ufeffin\ufeffthe\ufeff(O,\ufeff\u03b8,\ufeffz)\ufeffcoordinate\ufeffsystem.\ufeffFigure\ufeff1\ufeffillustrates\ufeffthe\ufeffschematization\ufeffof\ufeffplain\ufeff cylindrical\ufeffjournal\ufeffbearing.\nEquation Fluid Flow\n1.\ufeff\ufeff Equations of Continuity:\ufeffThis\ufeffequation\ufeffcan\ufeffbe\ufeffexpressed\ufeffby\ufeffthe\ufefffollowing\ufeffform\ufeff[10]:\n\u2207( ) =\u2192 \u03c1U 0 \ufeff (1)\nwhere:\nU\ufeff(u,\ufeffv,\ufeffw)\ufeffis\ufeffthe\ufeffvelocity\ufeffvector\ufeff \u03c9\ufeffis\ufeffangular\ufeffvelocity\ufeff \u03c6\ufeffis\ufeffattitude\ufeffangle\ufeff \u03b8 is\ufeffangular\ufeffcoordinate\ufeff", + "Equation\ufeff(1)\ufeffcan\ufeffalso\ufeffbe\ufeffwritten\ufeffas\ufefffollows:\n\u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 = u x v y w z 0 \ufeff (2)\n2.\ufeff\ufeff Navier-Stokes Equations:\ufeffThe\ufeffNavier-Stokes\ufeffequation\ufeffcan\ufeffbe\ufeffdefined\ufeffin\ufeffthe\ufefffollowing\ufeffform\ufeff(1997):\n\u2202\u2207 \u22c5 \u2297 = \u2212\u2207 + \u2207 \u2207 + \u2207 + \u2192 \u2192 \u2192 \u2192 ( ) .( ( ) )U U p U U BT\u00b5 \ufeff (3)\nwith,\ufeffP\ufeffstatic\ufeffpressure\ufeff(thermodynamic);\ufeffU\ufeffvelocity;\ufeff\u00b5\ufeffdynamic\ufeffviscosity. For\ufefffluids\ufeffin\ufeffa\ufeffrotating\ufeffframe\ufeffwith\ufeffconstant\ufeffangular\ufeffvelocity\ufeff\u03c9,\ufeffsource\ufeffterm\ufeffB\ufeffcan\ufeffbe\ufeffwritten\ufeff as\ufefffollows:\nB U r = \u2212 \u00d7 + \u00d7 \u00d7( ) \u2192 \u2192 \u2192 \u2192 \u2192\u03c1 \u03c9 \u03c9 \u03c92 \ufeff (4)\nEquation\ufeff(1)\ufeffcan\ufeffalso\ufeffbe\ufeffexpressed\ufeffin\ufeffthe\ufeffform:\n\u03c1 \u00b5( ) ( )u u x v u y w u z p x x u y u z u XB \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 = \u2212 \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 +\n2\n2\n2\n2\n2\n2 \ufeff\n\u03c1 \u00b5( ) ( )u v x v v y w v z p y x v y v z v YB \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 = \u2212 \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 +\n2\n2\n2\n2\n2\n2 \ufeff\n\u03c1 \u00b5( ) ( )u w x v w y w w z p z x w y w z w ZB \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 = \u2212 \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 + \u2202 \u2202 +\n2\n2\n2\n2\n2\n2 \ufeff\n\u03c1\ufeffis\ufefffluid\ufeffdensity. In\ufeffregard\ufeffto\ufeffthe\ufeffZ\ufeffaxis\ufeffas\ufeffthe\ufeffrotation\ufeffaxis,\ufeffof\ufeffthe\ufeffcomponents\ufeffB\ufeffmay\ufeffbe\ufeffexpressed\ufeffas\ufefffollows:\nx z x zB r v= +( )\u03c9 \u03c92 2 \ufeff\ny z y zB r u= +( )\u03c9 \u03c92 2 \ufeff (5)\nzB = 0 \ufeff\nThe\ufefffinite\ufeffvolume\ufeffmethod\ufeffused\ufefffor\ufeffsolving\ufeffthe\ufeffcontinuity\ufeffequations\ufeffof\ufeffNavier-Stokes\ufeff[9],\ufeffconsists\ufeff in\ufeffsubdividing\ufeffthe\ufeffphysical\ufeffdomain\ufeffof\ufeffthe\ufeffflow\ufeffvolume\ufeffelements\ufeffmore\ufeffor\ufeffless\ufeffregular,\ufeffshe\ufeffconverted\ufeff the\ufeffgeneral\ufeffdifferential\ufeffequation\ufeffsystem\ufeffalgebraic\ufeffequations\ufeffrelating\ufeffthe\ufeffvalues\ufeffof\ufeffthe\ufeffvariable\ufeffof\ufeffthe\ufeff adjacent\ufeffnodes\ufeffof\ufeffa\ufefftypical\ufeffvolume\ufeffcontrol.\ufeffThis\ufeffis\ufeffobtained\ufeffby\ufeffintegrating\ufeffthe\ufeffgoverning\ufeffdifferential\ufeff equation\ufeffin\ufeffthis\ufeffvolume\ufeffcontrol." + ] + }, + { + "image_filename": "designv11_80_0000601_ceit.2018.8751878-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000601_ceit.2018.8751878-Figure2-1.png", + "caption": "Fig. 2. Illustration of inertial and body frames.", + "texts": [ + " 1 arrow width indicates rotation speed of rotor where thicker arrow means faster rotating rotor with respect to others. Since quadrotor is a highly non-linear system, some assumptions are made to make the modeling easier. These are; 1- Body is symmetric and rigid. 2- Flapping and body deformations are negligible. 3- Center of mass coincides with the body frame origin. The quadrotor parameters used in this work and listed in Table I are taken from the example in [18]. Movement of a quadrotor can be analyzed in two different frames as the body frame and the inertial frame. In Fig. 2 the body frame is given the index \u201cG\u201d and the inertial frame is denoted with \u201cA\u201d. The resultant rotation transformation matrix \u201cR\u201d from body frame to inertial frame is given in (1) where \u201cs\u201d is the sine and \u201cc\u201d the cosine of related Euler angles. R= \u2212 \u0424 \u2212 \u0424\u0424 + \u0424\u0424 \u0424 + \u0424 \u0424 \u2212 \u0424 \u0424 (1) In (2) the position of the quadrotor in the inertial frame as (x, y, z) is defined with the vector \u201c\u03be\u201d and the angular positions form the Euler-angles vector \u201c\u03b7\u201d. In the body frame, the linear velocities form the vector \u201cVG\u201d and the angular velocities the vector \u201c\u03b3\u201d" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure16-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure16-1.png", + "caption": "Figure 16 Tenth order", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0000257_2019001-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000257_2019001-Figure11-1.png", + "caption": "Fig. 11. Overall deformation nephogram of sliding pin.", + "texts": [ + " However, in contrast to the differential pressure, the frequency variation induced deformation variation is relatively small, and the deformation growth rate is stable. Figure 10 shows sleeve stress nephogram when the frequency is 10Hz and a scatter diagram. As shown in Figure 10, the maximum sleeve stress increases with frequency and the range of the maximum stress gradually expands. In the range of 20\u201340Hz, the stress variation rate is moderate and stable; however, at 10Hz or 50Hz, the curve variation rate has an apparent significant change. This means that when frequency is too high or too low, stress change is apparent. Figure 11 shows sliding pin deformation nephogram when the frequency is 10Hz and a scatter diagram. As shown in Figure 11, the overall deformation of the sliding pin in the tripod sliding universal joint gradually increases with frequency; deformation at the ring bulge is greater than that of the sliding bar. When the frequency is low, deformation variation is small. As frequency increases, deformation increases significantly. When the frequency grows to a certain level, the deformation increment slows; however, the overall trend grows. The trend of deformation variation rate initially increases and subsequently decreases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000182_012029-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000182_012029-Figure2-1.png", + "caption": "Figure 2. Layout of the thermocouple (T1\u2013T6) and oil supply nozzles (N1\u2013N3).", + "texts": [ + " Monitoring the test modes, evaluating the technical condition of the examined bearings and the state of the equipment of the test bench was carried out by registering the following parameters: \u2022 rotational speed (n); \u2022 applied axial (Fa) and radial (Fr) loads; \u2022 temperatures of outer (T1-T3) and inner (T4-T6) ring of bearing; \u2022 oil temperature, inlet (Toil_in) and outlet (Toil_out); \u2022 oil flow rate (V); \u2022 vibrations. For lubrication and cooling, synthetic IPM-10 oil was used, with a viscosity of 3.47 cSt at 100\u00b0C. Filtration of the circulating oil was carried out with a 10-micron filter. The tests were carried out at different oil flow rates through the bearing (from 2 to 10 l\u00b7min-1). The amount of oil was determined by the number and diameter of the nozzles, as well as by the oil pressure in the oil system. The installation of the nozzles is shown in Figure 2. To ensure a flow rate in range from 2 to 5 l\u00b7min-1, one nozzle (N1) was installed; for a flow rate in range from 5 to 7 l\u00b7min-1, two nozzles (N1 and N2) were installed; and for a flow rate up to 10 l\u00b7min-1, three nozzles (N1, N2 and N3) were installed. To determine the temperature state of the bearing, thermocouples were installed on the inner and outer rings of the bearing. Thermocouples were also installed to measure the oil temperature at the inlet and outlet of the bearing. The layout of the thermocouples is shown in figure 2. A photograph of the bearing with the installed thermocouples is shown in figure 3. ToPME IOP Conf. Series: Materials Science and Engineering 489 (2019) 012029 IOP Publishing doi:10.1088/1757-899X/489/1/012029 To plan an experimental study of heat generation in split inner-ring ball bearings it is necessary to choose the sufficient number of experiments to obtain an empirical relationship, by which it is possible to determine the heat generation in a ball bearing under different operational conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003452_0954405420951101-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003452_0954405420951101-Figure6-1.png", + "caption": "Figure 6. The tooth engagement period.", + "texts": [ + " The maximum stress is calculated for different number of tooth and coefficient of bottom clearance. The calculation procedure of maximum root stress is shown in Figure 5. The maximum root stress of the TRT is studied by assuming a normal load at the Highest Point of Single Tooth Contact (HPSTC). The location of the HPSTC is calculated as the joint point between single tooth-meshing and double tooth-meshing. The single tooth-meshing section and the double tooth-meshing section in an engagement period are shown in Figure 6. The pressure angle on pitch circle, addendum circle of Gear 1 and addendum circle of Gear 2 are noted as a1, a2, aa1, and aa2, respectively. The contact ratio can be calculated as: e= z1 tanaa1 tana1\u00f0 \u00de+ z2( tanaa2 tana2)\u00bd =2p \u00f012\u00de The radius of the switching point between the single teeth-meshing area and the double teeth-meshing area can be calculated by: rj1= ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2b1+ ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r1 + ha1m\u00f0 \u00de2 r2b1 q (e 1)pmcosa1 2s \u00f013\u00de The pressure angle on this point can be obtained by: aj1 = arccos (rb1=rj1) \u00f014\u00de The most common mathematic model to calculate the maximum tooth root stress is 30 incline tangent method recommended by ISO,15 which is shown in Figure 7" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002698_b978-0-12-821918-8.00003-6-Figure3.102-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002698_b978-0-12-821918-8.00003-6-Figure3.102-1.png", + "caption": "Fig. 3.102 Schematic representation of RPM laser metal deposition system. (Reichardt, A., Dillon, R.P., Borgonia, J.P., Shapiro, A.A., McEnerney, B.W., Momose, T., Hosemann, P., 2016. Mater. Design 104, 404. With kind permission of Materials and Design.)", + "texts": [ + " Indirect information on the hardness in 304L stainless steel can be obtained from a functionally graded Ti6Al4V to 304L component with a vanadium interlayer fabricated by laser metal deposition additive 154 Additive and traditionally manufactured components manufacturing. In the process, a melt pool is formed by rastering the laser beam across the sample surface, and powder is injected into the melt pool to deposit each layer. In the case considered, namely functional composition gradient, dissimilar material is deposited by layering an alloy directly on a dissimilar material. The gradient components were additively manufactured using a four hopper RPM (RPM and associates, In.) laser deposition system shown schematically in Fig. 3.102. The Vickers hardness measurement along the length of the component of interest, namely SS 304L is shown in Fig. 3.103. The hardness through the pure component of 304L is quite constant at a level of 185 Hv. The Vickers hardness indicated in Fig. 3.103 was confirmed in an earlier work evaluating the mechanical properties in 304L stainless steel also fabricated by additive manufacturing. This can be seen in Table 3.28. Note, however, that the Vickers hardness is layer thickness-dependent, decreasing with layer thickness increase" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001575_s40684-019-00174-6-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001575_s40684-019-00174-6-Figure2-1.png", + "caption": "Fig. 2 Illustrations of a scanning strategy and b laid-up orientation and pull-out force in tensile test", + "texts": [ + "\u00a0(10) and the average energy density can be simplified as Eq.\u00a0(11): where dV is the product of dw , dh , dv and dt ; w is the hatch distance; h is the layer thickness; and v is the scanning speed. When the input heat was considered to be absorbed in the body and skin layers, the overall energy density will be the sum of Eqs.\u00a0(9) and (11). The average energy density can be expressed as Eq.\u00a0(12): All tests were performed at the additive manufacturing centre (Renishaw AM250, UK), with laser power and spot size up to 200\u00a0W and 75\u00a0\u00b5m respectively. Figure\u00a02a shows the layout for the features in SLM whilst Fig.\u00a02b illustrates the shape, laid-up orientation of the layer body and the applied pull-out force in the tensile test. In the selective laser melting process, scanning speed (250\u2013570\u00a0mm/s) was varied according to the function of the layers, such as inter-layer supports, fillers or exterior skins. Figure\u00a03a presents the setup of the selective laser melting which was performed on the Ti substrate (CP01, thickness: 0.8\u00a0mm) and layered with Ti-6Al-4V powders for geometrical features. Table\u00a02 lists the (8) repetition \u22c5 Plaser dV + [ x ( k T x ) + y ( k T y ) + z ( k T z )] = c T t " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000586_mi-12-2018-0084-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000586_mi-12-2018-0084-Figure1-1.png", + "caption": "Figure 1 Enzymatic electrode structure.", + "texts": [ + " To obtain graphene conductive paste for printing electrode layers, polymethyl methacrylate Mw 350000, diethylene glycol monobutyl ether acetate and GNPs, with 15-mm medium diameter, and up to 10nm of thickness were used. The biosensing substrates were made using the AChE enzyme (Type VI-S, lyophilized powder, electric eel source), acetylthiocholine chloride (ATChCl), tetraethyl orthosilicate (TEOS), hydrochloric acid (HCl; 37 per cent in H2O), Triton X-100 and phosphate buffer (PBS) pH 7.0. As an organophosphorus pesticide, Malathion PESTANAL\u00ae was chosen. Electrodes consisted of an electrode with enzymatic ink, an insulated conductive path and a contact, screen-printed on a flexible substrate (Figure 1). A PET film 100 mm in thickness was chosen as a support material for the electrodes to ensure flexibility, electrical resistance and water resistance. First, electrical paths with contacts were screen-printed on the PET foil using silver paste. After 30min of curing in 120\u00b0C, GR electrode pad was screenprinted with Electrodag 423SS onto the silver contact. To evaporate the excess of the solvents, the samples were dried at 120\u00b0C for 1 h. After drying, the dielectric paste (Electrodag 452SS) was screen-printed to insulate all parts except electrode and contact, and then cured in UV light for 15 s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002607_tec.2020.2989700-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002607_tec.2020.2989700-Figure6-1.png", + "caption": "Fig. 6. A finite element model of a LSIPM motor.", + "texts": [ + " For a given operating frequency of the motor \ud453 , the frequency ordering of the nodes at Level 5 is illustrated in Fig. 5. The frequency band of each node is \ud453 /4 Hz. The variations of energy in theses nodes are found to be most prominent during limit cycles induced operation and can be detected by observing the trajectories of their feature coefficients. Finite element (FE) analysis is conducted to investigate the limit cycle phenomenon using a three-phase 1-HP 208V LSIPM motor drive modeled by the software ANSYS Maxwell. Fig. 6 displays a FE model of the motor. The load torque is varied periodically to instigate the limit cycle phenomenon in the system. The load torque is created as time varying based on the following relationship: \ud447 (\ud461) = \ud434 + \ud435 sin(2\ud70b\ud453 \ud461) (27) where A and B are coefficients of the load torque, and \ud453 is the Authorized licensed use limited to: University of Exeter. Downloaded on June 22,2020 at 06:53:50 UTC from IEEE Xplore. Restrictions apply. 0885-8969 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001090_032034-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001090_032034-Figure1-1.png", + "caption": "Figure 1. Automatic gear-lining line a) draft machine and b) finishing gear reproducing satellite of rear axle differential, c) automated step conveyor d) working area of gear reprocessing machine before performing the operation.", + "texts": [ + " In the event that a contact patch of an inappropriate area and (or) location was initially obtained in a gear-over operation, subsequent operations can only worsen them or leave them Mechanical Science and Technology Update IOP Conf. Series: Journal of Physics: Conf. Series 1260 (2019) 032034 IOP Publishing doi:10.1088/1742-6596/1260/3/032034 unchanged. Therefore, it is precisely for the operations of draft and finishing gearing that it is important to ensure effective management of the manufacturing quality of the teeth of the satellite crown. In production, the process of toothprocessing is performed on machines ST 268 of the draft (Figure 1a) and finishing gearing (Figure 1b). They are rigidly connected with each other by an automated automated step conveyor (Figure 1.1 c) into an automatic line. During processing, the satellites move along the conveyor strictly in accordance with the beat. The capacity of the conveyor is 14 satellites. The satellite billet is installed by the auto operator into the expanding horizontal mandrel, the broach rotates around the vertical axis and receives a coordinated translational movement parallel to the line of the depressions of the gear wheel. Both machines are structurally similar, the difference lies in the performance of interchangeable sections of circular broach and copiers, defining the trajectory of their relative movement relative to the stretched workpiece" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002299_rteict42901.2018.9012428-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002299_rteict42901.2018.9012428-Figure1-1.png", + "caption": "Fig. 1. RMxprt model of Line Start Permanent Magnet Synchronous Motor The motor is provided with inbuilt squirrel cage windings in the rotor. The permanent magnets are embedded underneath the cage windings. The motor is self-starting.", + "texts": [ + " FEM involves solving the Maxwell\u2019s Equations at every sub-domain element and adding up the fields to solve for the entire machine [12]. LSPMSM is modeled using ANSYS-Maxwell FEA software tool. FEM is used for solving complex electromagnetic field problems, while considering the non-linearity of the domain of analysis. In this work, Maxwell is used to model and analyze the LSPMSM. RMxprt, short for Rotating Machine expert, is a machine model in which the user can insert the design parameters of the machine [9]. The RMxprt LSPMSM model is as shown in Fig. 1. Adjustable speed synchronous machine module in ANSYS is considered for modeling LSPMSM. Hence it is directly connected to the three-phase supply. The magnetic and electric characteristic of LSPMSM is analyzed by co-simulating the machine from ANSYS-Maxwell with ANSYS-Simplorer. Table I shows the specifications of the LSPMSM considered. The machine in ANSYS-Maxwell is powered by a three phase, 240 V, 50 Hz source. The circuit elements available in Simplorer are used for building the source. Also, the machine winding inductances and resistances are brought out as individual components in Simplorer as it is not possible to incorporate them in 2-D FEM model of the machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003964_j.ijleo.2020.165806-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003964_j.ijleo.2020.165806-Figure3-1.png", + "caption": "Fig. 3. 6-DOF parallel mechanism.", + "texts": [ + " Optik 226 (2021) 165806 the traditional development design process, making the entire design process achieve full digitalization and automation, and ultimately solving a series of the design optimization problems more rapidly and efficiently [22,23]. In summary, this paper attempts to apply this software to the parameter error identification. To verify the effectiveness of the above method, a parameter calibration test was performed on the parallel mechanism. In this paper, the object of parameter calibration is a set of 6-DOF 6-SPS parallel mechanisms, as shown in Fig. 3. Its structural parameters are shown in Table 1. The calibration test elements include 6 displacement sensors, sensor stand, and reference block to be measured (target reference block), etc., as shown in Fig. 4. The displacement sensors are SM30 series Czech ESSA grating displacement sensors, which have the advantages of large measurement range, high resolution and high accuracy, etc. [24]. Based on the 222-type ODM system, the calibration test was carried out, and the photo of the test site is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure33.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure33.4-1.png", + "caption": "Fig. 33.4 3D transient temperature distribution at different positions and different layers during five-layer laser additive manufacturing of Ti\u20136Al\u20134V alloy", + "texts": [ + " 1 2 3 4 5 6 Power (W) 100 300 200 200 300 400 Speed (mm s\u22121) 200 200 200 300 400 500 0.13 mm. Moreover, a percentage difference in melt-pool depth is estimated to be 5.38% for all processing conditions of Table 33.1 and which is under acceptable range. Comparison of computed (left) and experimental (right) melt-pool shape corresponding to data set #3 of Table 33.1 is presented in Fig. 33.3, and the red color zone represents the melt zone. With the application of laser power, a sudden rise in temperature is observed with a high-temperature gradient. Figure 33.4 shows the 3D transient temperature distribution at different positions and at different layers corresponding to process variables given in Table 33.1. Figure 33.4a shows the location of the laser source 33 FE-Based Heat Transfer Analysis of Laser \u2026 387 which is at the beginning of the first layer corresponding to data set #6. Figure 33.4 (b) represents the location of the laser source which is at a distance of 7.75 mm along the scanning direction of the second layer corresponding to data set #5. Similarly, Fig. 33.4c demonstrates the location of the laser source which is at the center of the third layer corresponding to the data set #4. Figure 33.4d represents the location of the laser source which is at a distance of 23.25 mm along the scanning direction of the fourth layer corresponding to data set #3. Figure 33.4e shows the location of the laser source at the end of the fifth layer corresponding to data set #2. Figure 33.4f represents the cooling nature of laser additive manufacturing of Ti\u20136Al\u20134V alloy. It is observed from Fig. 33.4c, d that the melt zone and heat-affected zone are wider for high laser beam power and low scanning velocity. Figure 33.5 depicts the time\u2013temperature history of each layer at a distance of 7.75 mm along the scanning direction. This figure shows a rapid rise in temperature followed by a gradual decrease in temperature. The sudden rise in temperature is due to sudden impingement of laser power at a point, and a gradual decrease is because of laser source moving away from that specific point. Theoretically, there should be five different peaks for the first layer, four for the second layer and three for the third layer and likewise" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003121_0954405420932445-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003121_0954405420932445-Figure10-1.png", + "caption": "Figure 10. Miniature bevel LG workpiece: (a) 3D model of workpiece. (b) Optical profiler picture of workpiece.", + "texts": [ + " Conclusions are drawn from above that by selecting the proper process parameters, the layer-by-layer NSPLA method is capable of manufacturing micro cones with semi-top angles within a range of \u00bd258, 708 . Then, an example of machining a miniature bevel LG with a suitable size using NSPLA is presented. The semi-top angle of the envelope cone of the bevel LG is 45 ; the number of teeth is 3, the thickness of the tooth profile is 20mm; the diameter of the bottom circle is 1375 mm; and the range of t is t 2 0:24p, 0:94p\u00bd . By using equations (4)\u2013(6), the parametric equation of the contact curve of the miniature bevel LG is derived as equation (5), in units of mm. Figure 10(a) shows the target workpiece of the miniature bevel LG R= x2 =329t cos (t) y2 =329t sin (t) z2 = 329t 8< : t 2 0:24p, 0:94p\u00bd \u00f05\u00de According to the fitting results of the experimental data, three combinations of scanning speed and laser power parameters can be selected for machining a cone with a semi-top angle of 45 as follows: 500mm/s, 42.89%; 750mm/s, 46.62%; and 1000mm/s, 49.81%. The three laser power parameters are relatively close, so all three combinations of the parameters are used in the manufacturing process", + " Step 2: Manufacturing of the teeth The basic requirements of manufacturing the LG teeth are as follows: to remove the material in the tooth space and to ensure the integrity and precision of the contact curve at the same time. The following parameters are set as the parameters of machining the teeth: 30% of the maximum laser power; scanning speed of 500mm/s; two-way reciprocating vertical scanning mode with a hatch of 5mm; scan times of 12. Three teeth with a thickness of 20mm were manufactured on the cone. Figure 10(a) is the 3D model of the target workpiece to be machined, and Figure 10(b) is the surface morphology of the machined workpiece measured by the optical profilometer. A comparison of Figure 10(a) and (b) confirms that the surface morphology of the workpiece is consistent with the expected morphology. The results of the shape error of the teeth profiles are shown in Figure 11. The three contact curves are shown as yellow, blue and red solid lines. The maximum profile deviation of the contact curve is fH=50mm. Figure 12 shows SEM images of the miniature bevel LG. Figure 12(a) shows that the overall external diameter of the sample is approximately 1.4mm, and the symmetry and integrity of the sample are good" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000020_mwscas.2018.8623843-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000020_mwscas.2018.8623843-Figure2-1.png", + "caption": "Fig. 2. Quadrotor with standard configurations", + "texts": [ + " (3) in Theorem 1, and dynamical system state equations of translations for the multirotor are written as the explicit form: d dt ( r r\u0307 ) = ( r\u0307 \u2212ge3 + 1 m B(\u03d51(t, (x0, x\u03070)T, u))Ftra(u) ) . (6) In this paper, we assume that all rotors are the same, distributed evenly and coplanar, and the distance from each rotor to the geometric center of the multirotor \u2113 is equal [3]. The multirotor have standard symmetrical configurations that the clockwise rotating rotor is adjacent to the counterclockwise rotating rotor as shown in Fig. 2\u20134. Fi, (i = 1, 2, . . . , 2p, p = 2, 3, 4) and Mi, (i = 1, 2, . . . , 2p, p = 2, 3, 4) of Figure 2\u20134, represent vertical force and moment, respectively. Each motor of multirotors has an angular speed \u03c9i and produces a vertical force Fi according to: Fi = kFi\u03c9 2 Mi, i = 1, 2, . . . , 2p, p = 2, 3, 4. (7) Each motor also produces a moment according to: Mi = kMi\u03c9 2 Mi, i = 1, 2, . . . , 2p, p = 2, 3, 4. (8) In practice, simple lumped parameter models are applied, such that kF > 0 and kM > 0 are given as constants that can be easily determined from static thrust tests. For multirotors from quadrotors: p = 2, hexarotors: p = 3 until octorotors: p = 4 that have standard symmetrical configurations, we define the moments of Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000505_14484846.2019.1626529-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000505_14484846.2019.1626529-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of the pocket-type orifice aerostatic bearing.", + "texts": [ + " In this paper, the pressure governing equations of orifice and porous aerostatic bearing were derived based on the gas-lubrication theory, then the dynamic and static characteristics of bearings with different restrictors were obtained with FLUENT software, and the superior bearing parameters were determined for the single-restrictor bearings. Based on the preceding optimization, the performance of two types of bearings with the different number of restrictors was investigated and compared. As a result, the performance characteristics and operating conditions of different restrictors were highlighted, which could effectively guide the optimization design of aerostatic bearings. 2.1. Orifice aerostatic bearings Figure 1 is a schematic diagram of the orifice aerostatic bearing with the single restrictor. The external compressed gas Ps feeds into the bearing clearance through the restrictor to form the gas film that can support the load, finally the gas flows into the ambient environment from the end of bearing. In order to obtain the pressure distribution in the gas film, the gas continuity equation, the gas-state equation and the simplified Navier Stokes equation are used simultaneously, certainly, the correct boundary conditions also should be introduced (Li and Ding 2012)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure13-1.png", + "caption": "Figure 13 Seventh mode", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0000864_012083-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000864_012083-Figure3-1.png", + "caption": "Figure 3. Normalization of the shape and relative position of the total contact patch a) nominal location, b) the allowable dimensions of the total contact patch, c) the maximum allowable boundaries of the location, d) the minimum allowable boundaries. Its relative location is normalized by establishing the minimum and maximum allowable areas for finding contact spots. The minimum allowable area is an area that limits the minimum allowable shape of the contact patch (Figure 3 g). The boundaries of the maximum allowable zone are set to prevent the total contact patch from escaping to the tooth boundaries under the maximum operating load (Figure 3c). The lines limiting the minimum and maximum areas of the total contact patch are equidistant to the corresponding opposite tooth profile. The position of the lines is normalized in the form of linear dimensions from the corresponding boundary of the zone of finding the total contact patch up to the borders of the gear tooth. They take values from 1 to 3 mm.", + "texts": [ + " The developed method is given additional modern requirements concerning its incorporation into the quality system of the enterprise [7], providing a collection of quantitative data on controlled characteristics, the identification of diagnostic components of the process of tooth processing [8], suitable for evaluating the result of processing as a result of mathematical modeling of the process [9]. As a result, ISTC-IETEM IOP Conf. Series: Materials Science and Engineering 570 (2019) 012083 IOP Publishing doi:10.1088/1757-899X/570/1/012083 to ensure the above requirements, in the developed method, the form of the total contact patch and its relative positioning are separately normalized. To establish the requirements of the contact patch shape, similar to the standard method, t the minimum and maximum limits of its boundary are set (Figure 3 b). In contrast with it, the nominal or ideal position of the total contact patch is given as an ellipse. For the considered bevel gear, according to the test results, its optimal position was found, shifted from the calculated transfer point towards the flat end (Figure 3 a). The minimum and maximum dimensions of the ellipse are set depending on the length of the tooth along the pitch circle and the average height of the tooth. Thus, the assessment of compliance of the total contact patch is simultaneously performed on a group of indicators: % L,% H - normalizing the shape of the total contact patch along the tooth and its height, as well as indicators L1, L2, L3, L4 - setting restrictions on the relative location of the contact spot. For each of these indicators, the upper and lower deviations are calculated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002029_iccas47443.2019.8971518-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002029_iccas47443.2019.8971518-Figure6-1.png", + "caption": "Fig. 6 Overall structure of drill bit and forelimbs.", + "texts": [ + " Using the rotating motor attached to the upper part, the rack moves forward and backward through the movement of the crank shaft and the worm gear, and the force is distributed and transmitted to the lower part having the same structure by using connecting the belt. In addition, a linear actuator and a link structure are designed to mimic the mole scapula and anatomy on both sides. Synthetically, the forelimbs move forward through one rotary motor and two linear actuators and make the motion that gathering paws to the center, pushing arms to both sides and returning to the initial position. The removal of excavated soil proceeds through the repetition of this motion. Figure 6 shows the overall design combined expandable drill bit and forelimbs. Drill bit module keeps advanced state while drilling, and when debris removal is started, it moves backward into the body, so that interference between the drill bit and forelimbs can be prevented. The sequence of the whole drilling mechanism of the proposed robot is as follows. 1) Advance drill bit module, 2) Rotate while expanding blades, 3) Excavation, 4) Contraction blades and reverse drill bit module, 5) Move forward and folding of forelimbs, 6) Remove debris spreading and move backward forelimbs", + " Using the rotating motor attached to the upper part, the rack moves forward and backward through the movement of the crank shaft and the worm gear, and the force is distributed and transmitted to the lower part having the same structure by using connecting the belt. In addition, a linear actuator and a link structure are designed to mimic the mole scapula and anatomy on both sides. Synthetically, the forelimbs move forward through one rotary motor and two linear actuators and make the motion that gathering paws to the center, pushing arms to both sides and returning to the initial position. The removal of excavated soil proceeds through the repetition of this motion. Fig. 5 Folding and spreading motion of forelimbs. Figure 6 shows the overall design combined expandable drill bit and forelimbs. Drill bit module keeps advanced state while drilling, and when debris removal is started, it moves backward into the body, so that interference between the drill bit and forelimbs can be prevented. Fig. 6 Overall structure of drill bit and forelimbs. The sequence of the whole drilling mechanism of the proposed robot is as follows. 1) Advance drill bit module, 2) Rotate while expanding blades, 3) Excavation, 4) Contraction blades and reverse drill bit module, 5) Move forward and folding of forelimbs, 6) Remove debris spreading and move backward forelimbs. Fig. 7 Drilling sequence of integrated system. During drilling, the drill bit is the part directly in contact with the ground, which is generated the most stress" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002475_asp-dac47756.2020.9045184-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002475_asp-dac47756.2020.9045184-Figure1-1.png", + "caption": "Fig. 1. Chiplet-based computer using an inductively coupled", + "texts": [ + "78-1-7281-4123-7/20/$31.00 \u00a92020 IEEE I. INTRODUCTION Embedded computer systems are becoming embedded into all our lives with the miniaturization and reduction of the power consumption. Among many applications, we are focusing on small systems that require complex mounting shapes such as micro-robots [1] and wearable devices using fibers [2]. In order to build a system with various shapes and robustness for such applications, we present a chiplet-based system utilizing wireless bus (Fig. 1). On-chip coils are formed along the outer periphery of each chip, and adjacent chips are wirelessly connected. This enables to construct embedded systems with various chip configurations, various shapes at low cost. On the proposed system, data transmitted by one chip is broadcast to all adjacent chips. As shown in Fig. 1, when a rectangular chip (Chiplet0) is used, it is expected that data is transmitted to all adjacent square chips (Chiplet1, 2, 3, and 4). Previously, no detailed investigation was conducted on the horizontal inductive coupling characteristics of rectangular on-chip coil. Although vertical bus communication between stacked chips has been investigated [3], horizontal bus communication has not been verified. In this paper, we show electromagnetic field simulation results of inductive coupling between rectangular coils" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001671_icems.2019.8921658-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001671_icems.2019.8921658-Figure4-1.png", + "caption": "Fig. 4. No load flux lines of the proposed BLDD machine. (a) Magnetic field produced by magnets on the inner rotor. (b) Magnetic field produced by stator magnets.", + "texts": [ + " TABLE I MAIN DESIGN PARAMETERS Parameter Regular BLDD machine Proposed BLDD machine Stator outer diameter 210mm 210mm Stator inner diameter 140mm 140mm Slot number 24 24 Modulation winding pole-pair 2 2 Regular winding pole-pair 11 5 Thickness of stator magnets -- 5mm Airgap length 1mm 1mm Flux modulator pole-pair 13 19 Thickness of flux modulator 5mm 5mm Inner rotor pole-pair 11 17 Thickness of inner magnets 3mm 3mm Stack length 50mm 50mm The no load flux lines of the proposed machine are shown in Fig. 4. To better introduce the effects of magnets on stator and inner rotor, their produced magnetic fields are presented respectively. When the magnets on the inner rotor are energized, 2-pole-pair field can be obviously observed in the stator. Similarly, 5-pole-pair field is produced by the stator magnets, which matches with the analysis based on flux modulation theory. The no load back EMF of the proposed machine is shown in Fig. 5. When \u03a9ro=500rpm and \u03a9ri=0, i.e., the inner rotor is static, the frequencies of back EMFs in two windings keep the same" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002274_j.promfg.2020.02.008-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002274_j.promfg.2020.02.008-Figure7-1.png", + "caption": "Fig. 7. Deposition tracks and appearances for a) block shape and b) thin wall.", + "texts": [ + " In the experiments with the coolant system, the same deposition conditions were employed with the experiments without cooling system. The appearance of experimental setup is shown in Fig. 6. For thermal observation, the same IRCamera was used again. The cooling system parameters given in Table 2 were kept constant for all experiments. In order to evaluate the mechanical properties of the deposited material, two types of shapes were produced and their scheme of the scan strategy can be seen in shown in Fig. 7. For the block shape, the deposition starts with the rectangular outline by depositing four lines as shown. Afterwards the inner part is deposited with a total of 22 lines. The deposition direction alternates with each layer. The deposition pitch was 1.5 mm, and the height of a single layer was 0.9 mm in the block deposition. For the thin wall, the scan strategy is simple since only a single line is deposited on the x-y plane, alternating the deposition direction with each layer. The height of a single layer was 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002493_042027-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002493_042027-Figure1-1.png", + "caption": "Figure 1 Finite element model", + "texts": [ + "The slewing bearing has a ball diameter of 40mm, a bearing outer diameter of 4980mm, a bearing inner diameter of 4880mm, an outer ring curvature radius factor of 0.53, and an inner ring. Coefficient of curvature radius 0.53, number of rollers 298. In order to facilitate the meshing and improve the quality of the mesh, the sliced solid is segmented. The upper and lower raceways use 20-node solid186 high-order hexahedral elements, and the rolling body adopts 8-node solid187. Tetrahedral element,as shown in Figure1. The plastic material model can not only reflect the elastic deformation behavior of the material, but also reflect the plastic deformation that occurs after the material reaches the yield limit. This material model simplifies the stress-strain curve into a polyline, where the yield strength\u03c3 It is 1280Mpa, Poisson's ratio 0.3, elastic modulus E is 210Gpa, and shear modulus Et is 2Gpa. The entire solving environment was performed in ANSYS Workbench. The contact algorithm selected extended Lagrangian multiplier algorithm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002550_icmre49073.2020.9064996-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002550_icmre49073.2020.9064996-Figure5-1.png", + "caption": "Figure 5. Propeller distribution of the robot.", + "texts": [ + " The center of gravity G of the robot is selected as the origin F of the moving coordinate system, \ud835\udc74\ud835\udc6d is only related to the structure distribution of the robot itself, as in \ud835\udc74\ud835\udc6d = [ \ud835\udc4011 \ud835\udc4012 \ud835\udc4021 \ud835\udc4022 ] = [ \ud835\udc5a\ud835\udc70\ud835\udfd1\u00d7\ud835\udfd1 \ud835\udc76\ud835\udfd1\u00d7\ud835\udfd1 \ud835\udc76\ud835\udfd1\u00d7\ud835\udfd1 \ud835\udc70\ud835\udc6d ] \ud835\udc3c\ud835\udc39 = [ \ud835\udc3c\ud835\udc65\ud835\udc53 \u2212\ud835\udc3c\ud835\udc65\ud835\udc53\ud835\udc66\ud835\udc53 \u2212\ud835\udc3c\ud835\udc65\ud835\udc53\ud835\udc67\ud835\udc53 \u2212\ud835\udc3c\ud835\udc66\ud835\udc53\ud835\udc65\ud835\udc53 \ud835\udc3c\ud835\udc66\ud835\udc53 \u2212\ud835\udc3c\ud835\udc66\ud835\udc53\ud835\udc67\ud835\udc53 \u2212\ud835\udc3c\ud835\udc67\ud835\udc53\ud835\udc65\ud835\udc53 \u2212\ud835\udc3c\ud835\udc67\ud835\udc53\ud835\udc66\ud835\udc53 \ud835\udc3c\ud835\udc67\ud835\udc53 ] \ud835\udc6a\ud835\udc6d(\ud835\udc97) is a Coriolis force matrix, which can be parameterized into a skew-symmetric matrix, satisfying the requirements of \ud835\udc6a\ud835\udc6d(\ud835\udc97) + \ud835\udc6a\ud835\udc6d(\ud835\udc97)\ud835\udc7b = \ud835\udc76, as in \ud835\udc6a\ud835\udc6d(\ud835\udc97) = [ \ud835\udc76\ud835\udfd1\u00d7\ud835\udfd1 \ud835\udc8e\ud835\udc68(\ud835\udc97\ud835\udfd0) \ud835\udc8e\ud835\udc68(\ud835\udc97\ud835\udfd0) \u2212\ud835\udc68(\ud835\udc70\ud835\udc6d\ud835\udc97\ud835\udfd0) ] \ud835\udc68(\ud835\udc82) = [ 0 \u2212\ud835\udc4e3 \ud835\udc4e2 \ud835\udc4e3 0 \u2212\ud835\udc4e1 \u2212\ud835\udc4e2 \ud835\udc4e1 0 ] \ud835\udc6d is the amount of external force and moment, including the restoring moment (\ud835\udc6d\ud835\udc6d), the rotating moment of the screw propeller (\ud835\udc6d\ud835\udc73) and the hydrodynamic moment (\ud835\udc6d\ud835\udc7a). Note \ud835\udc93\ud835\udc69 = [\ud835\udc65\ud835\udc35 , \ud835\udc66\ud835\udc35 , \ud835\udc67\ud835\udc35]\ud835\udc7b is the distance between the center of buoyancy and the center of gravity. In the case of gravity balance of buoyancy ( \ud835\udc6d\ud835\udc69 + \ud835\udc6d\ud835\udc7e = 0 ), \ud835\udc6d\ud835\udc6d can be obtained as in \ud835\udc6d\ud835\udc6d(\ud835\udf3c) = [ \ud835\udc45\ud835\udc38 \ud835\udc39(\ud835\udf3c\ud835\udfd0)(\ud835\udc6d\ud835\udc7e + \ud835\udc6d\ud835\udc69) \ud835\udc93\ud835\udc69 \u00d7 \ud835\udc45\ud835\udc38 \ud835\udc39(\ud835\udf3c\ud835\udfd0)\ud835\udc6d\ud835\udc69 ] = \ud835\udc5a\ud835\udc54 [ 0 0 0 \u2212\ud835\udc66\ud835\udc35\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf11\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf03 + \ud835\udc67\ud835\udc35\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf11\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf03 \ud835\udc65\ud835\udc35\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf11\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf03 + \ud835\udc67\ud835\udc35\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf03 \u2212\ud835\udc65\ud835\udc35\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf11\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf03 \u2212 \ud835\udc66\ud835\udc35\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf03 ] The propeller distribution in the horizontal direction is shown in Fig. 5, and the expression of F can be obtained as in (9). In the same way, we can also obtain forces in the vertical direction. \ud835\udc6d\ud835\udc73\ud835\udc99\ud835\udc9a = [ \ud835\udc4b\ud835\udc3f\ud835\udc65\ud835\udc66 \ud835\udc4c\ud835\udc3f\ud835\udc65\ud835\udc66 \ud835\udc4d\ud835\udc3f\ud835\udc65\ud835\udc66 \ud835\udc3e\ud835\udc3f\ud835\udc65\ud835\udc66 \ud835\udc40\ud835\udc3f\ud835\udc65\ud835\udc66 \ud835\udc41\ud835\udc3f\ud835\udc65\ud835\udc66 ] = [ (\ud835\udc39\ud835\udc3f3 + \ud835\udc39\ud835\udc3f4 \u2212 \ud835\udc39\ud835\udc3f1 \u2212 \ud835\udc39\ud835\udc3f2)\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udefe (\ud835\udc39\ud835\udc3f3 + \ud835\udc39\ud835\udc3f2 \u2212 \ud835\udc39\ud835\udc3f1 \u2212 \ud835\udc39\ud835\udc3f4)\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udefe 0 \u2212(\ud835\udc39\ud835\udc3f3 + \ud835\udc39\ud835\udc3f2 \u2212 \ud835\udc39\ud835\udc3f1 \u2212 \ud835\udc39\ud835\udc3f4)\ud835\udc670\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udefe (\ud835\udc39\ud835\udc3f3 + \ud835\udc39\ud835\udc3f4 \u2212 \ud835\udc39\ud835\udc3f1 \u2212 \ud835\udc39\ud835\udc3f2)\ud835\udc670\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udefe (\ud835\udc39\ud835\udc3f3 + \ud835\udc39\ud835\udc3f4 \u2212 \ud835\udc39\ud835\udc3f1 \u2212 \ud835\udc39\ud835\udc3f2)(\ud835\udc65\ud835\udc56\ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udefe + \ud835\udc66\ud835\udc56\ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udefe)] The hydrodynamic force can't be calculated simply by the formula. ANSYS is used to simulate the simplified model. TABLE I. VISCOUS HYDRODYNAMIC COEFFICIENT In the simulation, the RNG K-\u03b5 turbulence model and SIMPLE algorithm are used to solve the coupling field of velocity and pressure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003669_compe49325.2020.9200186-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003669_compe49325.2020.9200186-Figure1-1.png", + "caption": "Fig. 1. 3D view of Axial Flux PMSM", + "texts": [ + " The complete control strategy for AFPMSM drive system with PSO algorithm is presented in Section III. In Section IV, the simulation analysis is performed to validate the control strategy and to analyze the performance of motor drive system. Finally, conclusion is drawn in Section V. 676 Authorized licensed use limited to: SUNY AT STONY BROOK. Downloaded on October 03,2020 at 22:11:50 UTC from IEEE Xplore. Restrictions apply. II. MACHINE TOPOLOGY AND MATHEMATICAL MODEL The topology of 3 phase 12 pole Axial Flux PMSM is represented in Fig.1. The double air-gap dual rotor configuration used for design purpose consists of two identical rotors with a stator in between them and the rotors are held together by the shaft. So the configuration is also known as twin-rotor AFPMSM configuration. The three phase windings are wound on slotted stator core while permanent magnets are embedded on the rotor surface of each rotor core of the motor. Permanent magnets in each rotor are alternately magnetized along circumference direction with optimized trapezoidal shape to provide self-balance to the motor structure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003148_012030-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003148_012030-Figure9-1.png", + "caption": "Figure 9. AGV at the cross line", + "texts": [ + " Observe Figure 8, the AGV is placed on the line with 2 sensors on the left of the line and 2 sensors on the right of the line. 2nd 2019 ICERA Journal of Physics: Conference Series 1577 (2020) 012030 IOP Publishing doi:10.1088/1742-6596/1577/1/012030 To follow the line, the AGV makes use of only 2 sensors (R1 and L1). These sensors have a default value of 1 and changes to 0 when the sensor passes over the black line. The other 2 sensors are used as check point counter. The check points next to the black line are utilized by the AGV to determine its location compared to where it is going. Figure 9 highlights the 5 different check points that are next to the black line. The AGV receives commands from the masters over the wireless network in the form of a Byte. That byte is then converted into a number that the AGV uses to interpret the master\u2019s command. The possible commands from the masters are as followed: 0: stop all motions 1: Go from A to B 2: Go from A to C 3: Go from B to A 4: Go from C to A The AGV program is a case structured program in which each command from the master triggers a specific case that needs to be executed", + "1088/1742-6596/1577/1/012030 START Master\u2019s CMD = 0 Master\u2019s CMD = 1 Master\u2019s CMD = 2 Master\u2019s CMD = 3 Master\u2019s CMD = 4 EXE CASE 1 A to B YES EXE CASE 2 A to C YES EXE CASE 3 B to A YES EXE CASE 4 C to A YES STOP ALL MOVEMENTS YES GET CMD FROM MASTER NO CMD = COMMAND Figure 10. AGV program flowchart Each one of the cases shown in Error! Reference source not found., is a subroutine of its own that controls the AGV allowing it to follow a specific path. For example: in CASE 1, the AGV will follow the Path from A to B. Referring to Figure 9, it has been shown that there are 5 different check points. Note the case in which the AGV follow the path from A to C and using L2 to count the check points, observe that there will be a point where L2 will cross the line that goes to B and will count it as a check point as well. The actions perform by the AGV differs at different check points, for example: in case 1 the AGV must go from A to B. At the start, the check point count is equal to zero. Therefore, the AGV uses the sensors R1 and L1 to follow the line" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001683_icems.2019.8921881-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001683_icems.2019.8921881-Figure2-1.png", + "caption": "Fig. 2. Output characteristics of initial SRM. (a) Torque; (b) inductance.", + "texts": [ + " The torque ripple and the average torque with variable optimization parameters are calculated by 2-D FEM. Then, the multi-objective optimization is carried out by the PR models of the optimization objectives combined with the Pareto genetic algorithm (PGA). II. ANALYSIS OF INITIAL SRM Fig. 1 shows the topology of a 12/8 SRM, which has been taken as the initial motor to be studied. The SRM consists of stator, rotor, shaft and coils of stator winding. The key design parameters are listed in Table I. The torque and inductance profiles at 4000 rpm are obtained by FEM, as shown in Fig. 2. As can be seen from Fig. 2, for an ideal SRM, the inductance profile should be constant prior to the overlap of stator and rotor poles, so no torque is produced in this region. In fact, the inductance would nonlinearly increase prior to the overlap angle due to the fringing flux effect and local magnetic saturation, as presented in Fig. 3. The nonlinear variation of inductance leads to the unnecessary torque, which is one of the reasons for torque ripple. To quantitatively analyze the torque ripple of SRM, the torque ripple rate rT is defined as: 978-1-7281-3398-0/19/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure90.7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure90.7-1.png", + "caption": "Fig. 90.7 Distribution of Von Mises stress a and plastic strain b at first failure stage in liner", + "texts": [ + "32 MPa applied on inner surface of liner at plastic stage showed severe yielding with Von Mises stress of 309 MPa (refer Fig. 90.6a) and plastic strain of 0.76% (refer Fig. 90.6b) at the transition regions near to pole and equator region. Further, the first failure location of liner is seen at transition region near the 1080 R. Pramod et al. equator section. Failure is observed in terms of crack appearing in liner and complete rupture is observed as equator region gets completely detached from the liner section at Von Mises stress of 312 MPa (refer Fig. 90.7a) and plastic strain of 0.82% (refer Fig. 90.7b). From the analysis, it was also required to know, if the failure is occurring in weld zones of liner as weld specimen has lower strength (refer Table 90.2). Observation showed that second failure of liner in terms of crack propagated at the transition region in dome section at Von Mises stress of 358 MPa (refer Fig. 90.8a) and plastic strain of 0.99% (refer Fig. 90.9a). Also Figs. 90.8c, d and 90.9c, d shows the enlarged view of the failure regions with complete separation of ruptured parts. It is also noted from Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000841_iceee2019.2019.00065-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000841_iceee2019.2019.00065-Figure4-1.png", + "caption": "Figure 4. Flight controller loop diagram", + "texts": [ + " Accelerometer and gyroscope use SPI, barometer and magnetometer use I2C channels of microprocessor, while GPS and telemetry use UART channels. Voltage and current measurements are performed in ADC channel. The software consists of four fundamental operations: sensor data acquisition, filtering and estimation, controller calculations and motion command signal generation. In each software loop, these operations are repeated respectively. When software is considered as a loop, operations performed during a loop and frequencies of these operations are shown in Figure 4. Main control loop operates at a frequency of 400 Hertz and components with a frequency below 400 Hertz is processed in the main loop according to their frequencies. Operating frequency of the attitude controller has a significant influence on the stability of the flight control system as it directly generates motor motion command signals. Therefore, attitude controller frequency should be selected as high as possible. Because maximum input signal frequency of the ESC\u2019s is 430 Hertz, the attitude controller operating frequency is set to 400 Hertz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003120_s38311-020-0255-4-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003120_s38311-020-0255-4-Figure5-1.png", + "caption": "FIGURE 5 Prototype of the drive set swing arm (\u00a9 Trier University of Applied Sciences)", + "texts": [ + " This has several advantages: it enables torque vectoring when driving and when recuperating the vehicle, thus ensuring that the vehicle can be kept stable at moderate lateral accelerations even when the rear axle is braked electrically alone. This means that mechanical brakes on the rear wheels can be dispensed with completely and the recuperation potential can be expanded, thus further reducing energy power demand. The drive concept is realized by compact drive set swing arms, which as trailing arms with suspension struts simultaneously take over the wheel guidance and are a novelty in passenger cars, FIGURE 5. The very compact, water-cooled engines are located close to the axis of rotation of the drive set swing arm. This significantly reduces the unsprung masses compared to wheel hub motors, and the mechanical stress on the drive motors due to vertical acceleration of the wheels is lower. Torque and speed are transmitted to the drive wheels by a V-ribbed belt and components modified for this purpose from a standard P0-hybrid drive from the project partner Continental, as well as a constant planetary stage whose output shaft absorbs all operating forces from the respective rear wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001371_012040-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001371_012040-Figure7-1.png", + "caption": "Figure 7. Temperature field (t=0.1 s, s=0.1 m, \u2103). (a) Slicing map, (b) Wheel surface and (c) Rail surface.", + "texts": [ + "1088/1757-899X/657/1/012040 at the initial contact is respectively 11.03 mm and 23.15 mm. However, because of temperature rise, the contact area increases by 78.1% and the maximum pressure decreases by 42.6% at the end of sliding time (figure 6). 4.2. Temperature of wheel/rail friction pair The stripping and abrasion of wheel/rail surface are closely related to friction heat. Huge friction heat generates during the wheel sliding, and the heat will influence the stress of wheel/rail contact region (equations (4) and (5)). Figure 7(a) is temperature sectioning maps in longitudinal and transverse direction through the contact point. Figures 7(b) and 7(c) is temperature field of the wheel and rail surface respectively. The temperature field is similar to the figure 7(a) at every moment during wheel sliding. The contact location of wheel does not change, so the temperature on the wheel surface is much higher than that on the rail surface. The high temperature region locates on the wheel/rail surface. However, the influence area and depth of temperature are limited. For wheel and rail surfaces, figure 7 shows the highest temperature is respectively 1390\u2103 and 798.5\u2103. On the rail, the influence of temperature on the depth is no more than 3.28 mm. However, the temperature influence depth on wheel is two times as large as that on rail. In order to research the thermo-mechanical coupling effect, seven nodes are selected from rail surface and subsurface (figure 8(a)). The nodes locate at the depth of 0~14.92 mm below the rail surface. The relationship between node temperature and wheel sliding distance is presented in figure 8(b)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001358_uralcon.2019.8877690-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001358_uralcon.2019.8877690-Figure1-1.png", + "caption": "Fig. 1. The cross-sectional area of LIM with transverse magnetic flux:1 \u2212 inductor; 2 \u2212 inductor core; 3 \u2212 inductor winding;4 \u2212 conductive part of the secondary element; 5 \u2212 secondary element core", + "texts": [ + " Still it should be noted that the detailed study for the current density distribution in the winding of LIMTMF inductor have not been carried out yet. II. FORMULATING THE RESEARCH PROBLEM Linear induction motor with transverse magnetic flux is a relatively new type of LIM. Due to the features of its design LIMTMF is able to arrange high-speed movement of transport vehicles with transverse magnetic flux suspended in a magnetic field. It is necessary to consider a simplified motor design the shown in Fig. 1 to study the current in the frontal parts of a linear induction motor. Fig. 2 represents the layout of the current zones of LIM inductor with transverse magnetic flux (LIMTMF). In the middle part, under each inductor's pole we assume a continuous current density distribution, whose time and direction of motion varies according to the sinusoidal law (Fig. 2). In the given parts of the inductor, there is the only one component of the current density, directed perpendicular to the movement of the traveling magnetic field" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001154_978-3-030-29041-2_28-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001154_978-3-030-29041-2_28-Figure10-1.png", + "caption": "Fig. 10. Cross section profiles from the wear tracks on the bottom portion of the W2 component, for 3, 5 and 7 N of normal force.", + "texts": [ + " This approach also provides benefits in terms of feasibility since it combines conventional manufacturing up to the component\u2019s height where geometrical complexity begins with additive manufacturing, providing all the design freedom to comply with the final application. Costs are optimized through parallel processing, less building time and less raw material for the most expensive manufacturing process. The cross-section of the hybrid nozzle bushing and its application on an injection mould are shown on Fig. 10. AM Tooling for the Mouldmaking Industry 167 The SLM job was prepared on a ProX\u00ae DMP300 (3D Systems, Rock Hill, USA). The main equipment features are summarized on Table 1. The building job was setup on the equipment\u2019s processing software for a single build of each model to enable processing time evaluation. The evaluated models were the original nozzle bushing (ONB), the topologically optimized nozzle bushing (TONB) and the hybrid nozzle bushing (HNB). Figure 11 shows the job setup and building of the TONB model", + " Another point of interest resultant from the wear test, is the geometry of the wear track itself. Study of Laser Metal Deposition (LMD) as a Manufacturing Technique 235 With regard to the analysis of the profile of the wear track resulting on the flat specimen it can be verified that there is also a clear relationship between the transverse area of the cavity with the normal force exerted in the specimen during the test. It is possible to verify that with the increase of normal force there is also an increase in both the profile depth and the profile width (Fig. 10). The results of the upper and lower part of the specimen present a similar tendency and range, with higher values of depth and width of the cavity achieved for a normal force exerted in the specimen of 7 N. 236 F. Q. Ramalho et al. All of the wear tests conducted consisted of 1800 reciprocating cycles that ended up in a sliding distance of 7200 mm. In order to calculate the wear rate coefficient, the diameter of the sphere needs to be calculated, in order to achieve the depth and volume of the wear track" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001038_acc.2019.8814945-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001038_acc.2019.8814945-Figure4-1.png", + "caption": "Fig. 4. A representation of the factor graph used for probabilistic inference of CPG parameters. The model exploits conditional independence across constituents of the parameterization to infer a distribution of CPG parameters (P) using random variables that represent environment (E), behavior (B), and model (M).", + "texts": [ + " Using the insight made available by Distributed Correspondence Graphs [23] for approximating models with large symbolic representations, we assume conditional independence across constituents of the symbolic representation to make inference tractable, i.e., P\u2217 = arg max p1...pn\u2208P |N|\u220f i=1 p (\u03c6i = true|pi,B, E ,M) (7) As in [24], [23] we model the expressions for conditional probability distributions using log-linear models that learn features weights from a set of annotated examples (environment, behavior, model, and learned parameters) obtained from the genetic algorithm. A graphical representation of the factor graph is illustrated in Figure 4. We approximate the inferred distribution of CPG parameters from this model at run-time using beam search and use the most likely parameters to inform our sampling of candidate CPG parameters for performing the given task. This is done by modeling the conditional probability in Equation 7 as a learned function, i.e., P\u2217 = arg max p1...pn\u2208P |N|\u220f i=1 f (\u03c6i = true, pi,B, E ,M) (8) To evaluate the performance of the proposed model for inferring CPG parameters, we follow the procedure outlined in Figure 3 and vary parameters of the kinematic model and environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003452_0954405420951101-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003452_0954405420951101-Figure1-1.png", + "caption": "Figure 1. Illustration of the gear rolling system at SKLMT.10", + "texts": [ + " Section \u2018\u2018The mathematical model\u2019\u2019 explains the calculation method for stiffness and maximum root stress of TRT. Section \u2018\u2018The comparison of the maximum root stress\u2019\u2019 and section \u2018\u2018The impact factors on the stiffness of rolling tool\u2019\u2019 compare the maximum root stress and stiffness of standard and optimized profile of TRT with different parameters, respectively. Section \u2018\u2018Conclusion\u2019\u2019 concludes this paper. The geometry of the proposed gear tooth profile The structure of the gear rolling system with axialinfeed developed by SKLMT at Chongqing University is shown in Figure 1. Two rolling tools have the same rotational speed controlled by motors and transmission system, which consists of belt pulleys and worm gears. The racial-infeed cylinder adjusts the center distance according to different workpiece. The workpiece are pushed through the space between two rolling gears by axial-infeed cylinder and the workpiece is deformed into a gear with involute teeth through the meshing with rolling gears. The standard gear profile and optimized elliptical curve are shown in Figure 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000755_b978-0-08-102211-5.00004-8-Figure4.8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000755_b978-0-08-102211-5.00004-8-Figure4.8-1.png", + "caption": "Figure 4.8 Schematic illustration of roving frame.", + "texts": [ + " Drawing is an important process to improve the evenness of the yarn and it is normally performed twice in the spinning mill. 99Auxetic fibres and yarns For many spinning techniques, the sliver produced by draw frame can be spun into yarn. However, for ring spinning, sliver cannot be directly spun to yarn in the factory production and roving as a feeding material is required. The roving process draughts a thick sliver to become a thin sliver with some twists (roving) and builds an even package on the roving bobbin for the ring spinning feeding. The roving process is schematically shown in Fig. 4.8. It can be seen that the sliver is first taken from the can and is draughted by rollers and aprons to form a thin roving, and then the roving is twisted and wound on the bobbin by flyer and spindle. Finally, the ring-spun yarn can be produced from roving by the ring spinning frame. The ring spinning frame almost has the same function as the roving frame. It utilises roller draughting system to attenuate the roving and ring and spindle system to add twists on the yarn. The ring spinning system produces yarns with the best quality among all the manufacturing systems and it is versatile" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003785_etfa46521.2020.9212115-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003785_etfa46521.2020.9212115-Figure1-1.png", + "caption": "Fig. 1: The position of the four sensors on the robot.", + "texts": [ + "), the localization sub system can calculate the coordinates (in the robot-centric system) of the source in 3d space. This removes the requirement for a universal localization system. Additionally, our approach requires sensors that are easily accessible (light, radio, Bluetooth etc) as long as they return the intensity of the received signal. 1) Basic Scheme: The main idea is to use the RSSI measurements of four sensors in order to calculate the position of the electromagnetic source in the 3d space. These sensors (S0, SR, SL, SF ) should be placed on the perimeter of a circle with radius r as seen on Fig. 1. Each sensor returns a measurement E, defined by (1), which is proportional to the intensity of the source (W ) and the angle of incidence (\u03c6 ) and inversely proportional to the square of the distance between the sensor and the source. E = W 4\u03c0d2 cos\u03c6 (1) The triangle formed by the sensor, the source and the projection of the source onto the sensors plane can be seen on Fig. 2. The cosine of the angle of incidence is equal to the 683 Authorized licensed use limited to: Carleton University. Downloaded on November 03,2020 at 11:55:26 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure77.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure77.1-1.png", + "caption": "Fig. 77.1 Schematic of gear test rig", + "texts": [ + " In the present preliminary study, the influence of hardener to resin ratio of adhesive on the performance of adhesive gear is investigated with actual and repaired gears. Commercially available structural adhesive was used as a polymer matrix in this study. The adhesive system is a combination of two parts, namely resin (bisphenol A epichlorohydrin) and hardener (polyamidoamine). The suitable mixing ratio of the adhesive system is of 1:0.8 by weight as per supplier\u2019s instruction. Gear Test Rig. Figure 77.1 shows the schematic of the test rig for evaluating the adhesive spur gear performance. It has a 0.5 HP DC shunt motor with maximum speed up to 1500 rpm. It is coupled with an alternator having the same capacity. The input to the motor and the field of the alternator is controlled externally by two separate autotransformers. The autotransformer is used to adjust speed and torque. There is an additional rectifier circuit consisting of diodes and capacitor for the AC to DC conversion. A 100 X rheostat is used as the load across the alternator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000029_s12239-019-0013-z-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000029_s12239-019-0013-z-Figure4-1.png", + "caption": "Figure 4. Double wishbone type (RSSR-SS) suspension mechanism with wrenches acting on the wheel hub.", + "texts": [ + " Again, the intersecting point of the instant screw axis and the plane of motion is the instant center of velocity of the coupler with respect to the ground. In this section, a method to find the twist of the wheel of a three-dimensional suspension mechanism is proposed. For this, five independent wrenches acting on the wheel hub by the connected links for the bump-rebound motion and steering motion are determined, then the instant screw axis and pitch of the wheel for each motion can be obtained by the twist reciprocal to the five wrenches using Equation (9). 5.1. Instant Screw Axis of Bump-rebound Motion Figure 4 shows an RSSR-SS spatial mechanism, which is the kinematic representation of the double wishbone type suspension. The wheel hub is connected to the vehicle body by two R-S links, each of which exerts two zero-pitch force wrenches on the wheel hub acting at the center of the S joint as shown in Figure 1 (c), and by an S-S link which has a zero-pitch force wrench acts through the centers of two S joints as shown in Figure 1 (e). These five wrenches are independent, and the twist reciprocal to the five wrenches determines the instant screw axis of the wheel with respect to the vehicle body for the bump-rebound motion and its pitch", + " For the McPherson strut type suspension, the steering axis is defined as the line passing through the centers of the S joints of the S-C link and R-S link, the line joining b0 and a1 in Figure 10 (b). In the case of the 5-SS multi-link type suspension mechanism, however, the steering axis cannot be determined geometrically as in the above cases. Kinematically, the steering axis can be defined as the instant screw axis of the wheel hub with respect to the vehicle body during its steering motion. When the suspension is steered, the wrench W3 acting along the tie rod, c0c1, in Figure 4 to Figure 6 disappears because c0, the spherical joint of the tie rod on the vehicle body, is moved by other link as shown in Figure 7, and there are only four wrenches acting on the wheel hub. The reciprocal screws to these four wrenches form a second-order screw system. This reflects the fact that the front suspension mechanism has two degrees of freedom. In order to find the steering axis for the steering motion only, it can be assumed that the wheel center remains the same height in vertical direction during the steering motion (Lee and Ahn, 1993)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003453_022004-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003453_022004-Figure1-1.png", + "caption": "Figure 1. Devices for seeding by scattering way.", + "texts": [ + " The applied ordinary methods of sowing do not fully meet the characteristics of mountain and piedmont agriculture and do not provide the best conditions for realizing the potential productivity of mountain forage land. AGRITECH-III-2020 IOP Conf. Series: Earth and Environmental Science 548 (2020) 022004 IOP Publishing doi:10.1088/1755-1315/548/2/022004 The optimal method of sowing crops is a uniform distribution of seeds over the area of nutrition. Such conditions are ensured by seeding using a disc sowing apparatus. The proposed device [8-10] (figure 1, a) for seeding grasses consists of a frame 1, support wheels 2 and a technological tank 3 for placing seeds and fertilizers. In the technological tank 3 there is a mechanism 4 for feeding and dosing in accordance with the norms of seeds and fertilizers. The working bodies are centrifugal disks 5, equipped with blades 6. The drive of the working bodies is carried out from hydraulic motors 7 and 8. The fertilizer 9 of the seed and fertilizer flow is placed on top of the centrifugal disks 5 and is fixed. In the upper part, it has a height-adjustable emphasis 10, made in the form of a roller. The position of the stop 10 in height is regulated by a screw mechanism 11. The profile track 12 rests on the limit stop 10, which is rigidly fixed to the metering shutter 13 (figure 1, b), which is installed in the guides 14 and has the ability to move vertically. The elastic elements 15 press it against the limit stop 10. The case of the first hydraulic motor 8 is rigidly mounted on the frame 1, and the case of the second hydraulic motor 7 is at one end with the help of the axis 16 to the frame 1 with the possibility of rotation, and at the other end by a link 17 is pivotally connected to the lever 18, which in turn is pivotally connected to the rod 19 of the hydraulic cylinder 20, connected through oil lines 21 to the spool valve 22 of the hydraulic system of the device", + " Spool valve stem 22, by means of roller 23, is supported by cam 24, which is rigidly fixed to slider 25, which passes with a gap through the hole of thrust ring 26 fixedly mounted on the frame 1. Spring 27 is installed between thrust ring 26 and straight cam 24. At the end of slider 25 wheel 28 is fixed, which rolls over the surface of cam 29. Cam 29 is fixed on pusher 30, at one end of which a copy wheel 31 is mounted pivotally, having constant contact with the soil surface I spring 32 disposed between support washer 33 fixed to plunger 30 and thrust ring 34 fixed on frame 1. In the process of operation of the device on a flat area (figure 1, a), material from the technological tank 3 is fed to the centrifugal disks 5 with blades 6 using feed mechanism 4 and scattered over the soil surface. In this case, rod 19 (figure 1, b) of hydraulic cylinder 20 occupies the lowest position, holding lever 18 and rod 17 in the position in which the axis of the shafts of hydraulic motors 7 and 8 are located vertically. Spool valve stem 22 in this case is in its lowest position, which corresponds to the maximum stroke of slider 25 to the right. Needle guide 9 of the material flow above the centrifugal disks 5 is installed in the middle position, as a result of which the same amount of material is supplied to the working bodies", + " Since guide 9 is installed in the middle position, limit stop 10 is located in the middle of profile track 12. Metering valve AGRITECH-III-2020 IOP Conf. Series: Earth and Environmental Science 548 (2020) 022004 IOP Publishing doi:10.1088/1755-1315/548/2/022004 13 is located in the lower position due to elastic elements 15. The rate of application of grass and fertilizer seeds is determined by the position of limit stop 10, adjustable by screw mechanism 11. When the device is operating on a slope section (figure 1, b), copying wheel 31 with pusher 30 with support washer 33 and cam 29 fixed on it moves upward, which helps compress spring 32, and wheel 28 rolls over the lifting surface of cam 29, moving slider 25 to the left. However, spring 27 is compressed and wheel 23 is rolled over the surface of cam 24 to the axis of slider 25, and the stem of spool valve 22 moves up. As a result, oil from spool valve housing 22 under pressure enters hydraulic cylinder 20, which leads to the retraction of rod 19, as a result of which hydraulic motor 7 with centrifugal disk 5 through lever 18 and rod 17 is rotated by an angle determined by the steepness of the slope" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003148_012030-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003148_012030-Figure12-1.png", + "caption": "Figure 12. AGV side view", + "texts": [ + " The AGV, in this test, is placed in an environment where it has to follow a black line on a white surface. The purpose of this test is to check the line following code that is loaded on the AGV\u2019s motherboard to see if it will be able to correctly see and follow the line. The test is conducted as followed: Place the AGV on the line with the sensors on either side of the line. Turn on the switch. Then wait about 10 to 30 seconds for the program in the memory to initiate. Turn on the motors switch to start running the AGV Figure 12 shows the worker\u2019s side by side view. It serves as reference for the location of the switches that control the LabVIEW robotic starter kit. This test will also determine the optimum speed at which the AGV can properly follow the line without going off track This test is designed to check the communication between the AGV and the computer. A simple application runs on the PC to send commands to the AGV to tell it where to go. With each command received, the AGV performs a specific section of its control code" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000739_kem.813.261-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000739_kem.813.261-Figure4-1.png", + "caption": "Fig. 4. Load cell reference system", + "texts": [ + " The rubber parallelepipeds are glued to a specimen holder by means of an ethyl-cyanoacrylate glue being careful that the specimen leading edge is aligned to the specimen holder one, in order to guarantee the identical geometrical conditions for each test. The properties of the specimen have been obtained by means of dynamic mechanic analysis (DMA) test, resulting in diagrams of the storage modulus and tan( ) versus frequency, as shown in Fig. 3. Before starting the test campaign, the load cell should be calibrated in order to remove the eventual offsets. The calibration procedure consists of two steps. In the first step, the road is not placed in the tank and the arm is positioned horizontally to calibrate the Fz (Fig. 4), in order to allow the load cell to measure all the forces along the x direction, according to the local reference system in Fig. 4. In the second step the arm is positioned vertically to calibrate the Fx (Fig. 4). The load cell therefore measures all the forces in the z direction of the local reference system. Then, the road is positioned in the tank and the sliding length is set equal to the nominal value of 50 mm by the endless screw. During the sliding, the tread specimens deposited rubber particles on the road, due to the friction. Furthermore, the wear of the tread specimens and the \u201crubberization\u201d of the road were observed. Such phenomena should be taken into account in order to keep both the boundary condition and the working parameters as constant as possible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001404_pgsret.2019.8882707-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001404_pgsret.2019.8882707-Figure1-1.png", + "caption": "Fig. 1. The structure of designed motor.", + "texts": [], + "surrounding_texts": [ + "978-1-7281-2301-1/19/$31.00 \u00a92019 IEEE\nIn this study, we proposed an induction motor (IM) with squirrel cage by optimizing rotor slot structure for higher starting torque and efficiency. The physical parameters of the rotor were defined as variables by parametric approximation method. Totally, 19008 of analyzes were performed at each point within the specified limits. Changes of the efficiency, weight, nominal and starting torque were obtained to determine optimized induction motor.\nKeywords\u2014motor design, electrical vehicle, induction motor, lower cage, upper cage\nI. INTRODUCTION\nThe consumption of fossil fuels and the increasing number of vehicles are known to cause environmental pollution and climate change. The popularity and use of electric vehicles (EV), which are considered as an alternative solution to these problems, is increasing day by day. In electric vehicles, the drive system consists of electric motor, power electronics and control systems. In electric vehicles, the performance of the vehicle is directly related to the choice of the motor, as the torque transmitted to the wheels is produced by the electric motor. Therefore, the selection of an electric motor to meet the performance expectations for the vehicle is of great importance [1-3].\nThe motors used in EVs are generally composed of DC motors, switching reluctance motors, induction and synchronous motors. Some electric vehicle manufacturers are reluctant to use permanent magnet motors due to high magnet costs, rapid fluctuations in raw material prices and the negative effects of high heat on the magnet. Moreover, they decide not to use these engines in new generation vehicles. Due to the ease of control, the preferred DC motors cause frequent maintenance due to their commutator and brush structure. Torque ripple of switching reluctance motors\nhas caused these motors not to become as widespread as expected [2-6].\nToday, developments in power electronics technology have facilitated the control of AC motors. [1,2]. Induction motors come to the forefront with features such as easy production, low cost and frequent maintenance. [3-7]. In a study which examined the effect of rotor slots on the performance of the engine, two different induction motor were designed with open and closed rotor and 2 dimensional finite element analyzes of the designs were made. As a result of the study, it was found that the open slot structure increased the harmonic components in the air gap. [8]. In another study using similar method, the rotor of the squirrel cage induction motor was designed and analyzed in 10 different structures. As a result of the analyzes, the best performance was obtained by round type rotor slot structure. [9].\nIn the systems, the characteristic of the electric motor must be compatible with the application characteristics. For example; If an electric motor with the correct operating characteristics is not selected even for a fan application, problems such as low efficiency or poor performance may be encountered. Similarly, considering the driving profile of an EV to be used in the city, it is known that the stop-and-go cycle will be high and the urban speed limits will be low. So, the efficiency of the motor and in particular the take-off torque must be optimized for the city driving profile.\nIn this study, the optimization of the rotor slots of the previously designed induction motor [10,11] was realized by considering urban driving profile. By determining the variables related to the rotor slots to ensure the targeted engine performance, the efficiency, take-off torque, nominal torque and total weight of the motor were investigated.\nII. URBAN USE ELECTRICAL VEHICLE\nThe urban car use is very popular than the highway car use. Therefore, an alternative to buyers needs to be offered, if the use of non-environmentally friendly vehicles is not desired. In that case, a small, light and safe electric vehicle is an important alternative to reduce the release of harmful gases in the city. Hence, clean, safe and energy efficient urban transport system can be implement [12,13].", + "The goal of this study is to design an induction motor for ergonomic, energy efficient and safe electric vehicle in urban use. The energy consumption can be decreased with the contribution of optimized weight and well-designed an electric motor. As a result of this, the dependency on oil will be reduced with emission of greenhouse gas and pollution of noise and air [12,13]. It should not be forgotten, the percentage 40% of CO2 emission of road transport is caused by the urban mobility. In addition, the transport activities is reason of quarter of total harmful emission [14].\nThe advantages of urban electric vehicle can be summarized as follows: reduction of emission for clean air, urban quality due to health and life, cheap transportation. Furthermore, the urban use will not affected by some adverse situations such as vehicle range, battery capacity and cost and vehicle mass like highway transportation. Because, the proposed model will be designed for two passengers to reduce the mass of vehicle. To capture the advantages of electric vehicle some operations need to be done such as charging infrastructure and urban policies. Perhaps the greatest challenge will be the lack of charging stations for electric vehicles in urban use [13,15].\nFor these reasons, the design of the electric motor, which is considered as most essential component of electric vehicles, was focused. In this study, an efficient induction motor design was aimed by taking into consideration the above-mentioned points, considering urban use conditions and higher efficiency in the city.\nIII. DESIGNED INDUCTION MOTOR\nThe design and nominal power parameters of the induction motor, which was designed previously [3,4], are given in Table 1.\nIV. OPTIMAL ROTOR SLOT STRUCTURE\nThe performance of the electric motors depends on the physical properties that make up the structure and the characteristics of the materials used. As the variations in the rotor slot structure of induction motors directly affect magnetic and electrical circuit models, it also changes the motor performance. As known, the rotor characteristics of the induction motors with squirrel cage cannot change after production. Particularly, while designing of the squirrel-cage induction motor, the working characteristics should be considered and the criteria for optimum rotor structure, such as combinations of stator-rotor grooves, single cage or double cage, choice of rotor conductor type (aluminum or copper) etc., should be determined [5,16].\nIn the design, various motor manufacturers' data was reviewed and as a result of optimization studies, stator and rotor slot numbers were determined as 36 and 46 respectively. Also, double cage and single cage structures were investigated for rotor slot structure. The double cage structure rotor type was determined more suitable for application requirements in order to increase take-off and nominal working performances. the rotor cage, consists of aluminum and copper conductors, were analyzed with the same slot cross-sectional areas for equal comparison. As a result of the analyzes, the motor with aluminum conductors rotor is better about take-off performance [10,11].\nIn this study, the effects of changes in slot geometries on engine efficiency and take-off torque are analyzed to determine optimum rotor structure. The parameters representing the lower cage and upper cage of the double cage rotor are given in Figure 2. The slot structure given in Figure 2 was used in both the upper and lower cage of the rotor slots. Bs0 and Hs0 parameters, seen on Fig. 2(b), are representing the junction area of lower cage and upper cage. The variables that precise the lower cage and upper cage of the rotor and the initial values defined for these variables are seen on Table 2. Some of these variables were expressed as constant values considering design limitations", + "The parametric solution method used in the optimization of electrical machines is a practical, fast and reliable method. In this method, the parameters given in Table 2 are defined as variables in the analytical solution. Then, the solution range and the solution steps to solve the problem will be determined at the target accuracy. The number of solution steps directly affects the sensitivity of the analysis [17]. The limits of the parameters determined in the study are given in Eq. 1.\nAs a result of the analysis, the change in the motor efficiency according to the determined parameters is given in Figure 3. In order to examine each parameter change, one of\nthe four parameters was changed, while the other 3 parameters were as fixed. The data selected as constant was given separately in each graph.\nmmstepmmDmm\nmmstepmmCmm\nmmstepmmBmm\nmmstepmmAmm\n5.0,2.52.1\n2.0,6.21\n2.0,41\n1,154 =\u2264\u2264 =\u2264\u2264 =\u2264\u2264 =\u2264\u2264\n(1)" + ] + }, + { + "image_filename": "designv11_80_0000630_s10704-019-00367-9-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000630_s10704-019-00367-9-Figure3-1.png", + "caption": "Fig. 3 Uniaxial tension test\u2014finite element mesh and drawing (Kramer et al. 2018) with dimensions in mm", + "texts": [ + "8GPa obtained from this fitting process was significantly lower than the expectedYoung\u2019sModulus for AM316Lmeasured byWu et al. (2014) of between 173 and 190GPa, and also much lower than their reported value of 193GPa for wrought 316L. In the first series of simulations, the SFC3 uniaxial tension experiments (Kramer et al. 2018) were simulated to obtain the material parameters listed in Table 1. A drawing of the uniaxial tension sample (dimension in mm) and the corresponding finite element meshes are shown in Fig. 3. The original material characterization simulations and blind predictions for the challenge geometrywere all performed usingmesheswith simple four-noded, tetrahedral elements and a typical element edge lengthof 0.1mmin refined areas. In these transient dynamic simulations with mass scaling, displacements of all nodes gripped by the fixture were prescribed. In one set of simulations, equivalent plastic strain is used as the measure of damage (i.e. \u03b1\u0302 and \u03b2\u0302 equal to zero) and in the second set of simulations non-zero values for \u03b1\u0302 and \u03b2\u0302 were used" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure1-1.png", + "caption": "Fig. 1. Original model", + "texts": [ + " Some of the well-known CT reduction techniques such as such to shift or segment of the magnet pole [1]-[3], to optimize the magnet pole arc [2],[3],[4],[7] to skew or step skew in magnet or stator core [3],[8],[9], to reduce the magnet strength [1],[7] to vary the pole arc length and offset [1],[10],[11], to employ dummy slot in stator core [11]-[12], to shape magnet edge[12]. However, the most effective technique to reduce the CT in PMMs is employing the two steps of slotting the magnet edge. By slotting the two steps of slotting in the magnet edge, it could promise to provide a new flux barrier in the magnet surface and optimize the magnet flux distribution in the air gap of the PMMs. II. PMM MODEL PROPOSED The Initial model, one of step model and the two steps model of PMMs studied in this paper as illustrated in Figures 1, 2, and 3, respectively. In Figure 1, it observed that the magnet structure of the PMMs adopted the conventional model. In the model, the height of the magnet was homogenous in all parts of the magnet structure. Thus, the height in the edge of the magnet is the same as the height in the center. Figure 2 shows the one step slot employed in the magnet edge. By using one step slot causes the magnetic flux distribution to become changing and decreasing in the edge of the magnet. Also, the total magnetic flux flowing into the air gap becomes reduced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000542_978-3-030-20131-9_261-Figure109-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000542_978-3-030-20131-9_261-Figure109-1.png", + "caption": "Fig. 109. Bead string state diagram of touching the fixed travel switch.", + "texts": [], + "surrounding_texts": [ + "In chapter 2, three methods of automatic knitting of bead mat are put forward. Among them, Warp and Weft Automatic Weaving Method is the most simple, but the cost is higher and its stability is poor. The lock stitch sewing weaving method is rea lly good, but there exists three difficulties applied in the machine: The cross-sectional size of the hook must be smaller than the size of the bead hole as the hook needle needs to completely pass through the bead, which makes it difficult to hook the string smoothly. It is difficult to guarantee the parallelism between the cross-section of the loop ring and the end of the bead of the first row in steps 4-6. When weaving a larger size beaded cooling pad, threading the string into the longer aligned transverse bead holes becomes hard. So compared with the other two methods, single-line straight-through method, which has good stability and high knitting efficiency and is easy to be realized on the machine, is the focus of the discussion below. Based on this method, an automatic weaving device capable of weaving a beaded cool pad is proposed and designed. 4.1 Feeding Device Design Before weaving the beaded cooling pad, since all the beading arrangements are disordered, it is necessary to design a device that puts the beads into the weaving state in an orderly manner, which is called feeding device. Referring to the feeding mechanism of the firecracker weaving [14-16], Fig.14 shows the schematic diagram of the designed bead feeding mechanism device. Before the device runs, all the beads are placed in the hopper and the two guiding wheels. A small number of longitudinally are placed in a horizontal arrangement. When the two guide wheels rotate in opposite directions, the beads in the hopper will be putted into the guide groove and conveyed to the front of the beading device in an orderly manner. In order to avoid a rigid collision between the feeding device and the ball transported device, the end of the guiding groove is made by a material with better elasticity. 4.2 Design and Working Principle of the Beaded Pad Weaving Device Fig. 15 demonstrated the beaded pad weaving device. Since the figure is only for explaining the movement process of the beaded pad weaving device, the feeding device is not shown in the figure. And there are 7 motors in this device. Control motor A controls the movement of the threading device. Control motor B controls the movement of the movable line of the downlink line. Control motor C controls the movement of the movable stroke switch. Control motor D controls the rotation of the output port and the braided port. Linear motor E S. Ouyang et al.2544 drives the up-line feed ball push block movement. Linear motor F pushes the braided beaded cool pad unit into the braided port. Linear motor G drives the linear motion of the downlink line feed bead block. Before the device is operated, the downlink threading is first performed. After the downlink threading is finished, the string is installed into the beaded pad weaving device. The downlink line with heavy beads at the end is wrapped around the fixed pulley mounted on the frame. Then it passes through the downlink line movable seat, the through hole, the downlink end sleeve, the bead and the braided port successively, to reach the uplink line. And the uplink line is directly connected to the needle of the threading device, the first string of beads are moved to the corresponding position on the weaving port, as shown in Fig.16. After the string installation is completed, the motor A is manually controlled to make the driving roller be located between the two trapezoidal blocks on the movable seat rail of the threading device. The manually controlled linear motor E is to drive the uplink line to send th e beads push block moves, which is external bead conveyed from the feeding device. It causes the holes axis of the bead to coincide with the needle axis of the bead threading device and the up-line bead push block is in the beading state. The specific work ing process of the device is as follows: Method Research and Mechanism Design of Automatic Weaving\u2026 2545 Step 1: Under the drive of the control motor A, the bead threading device moves to the right. When the driving roller moves in the second trapezoidal block on the movable seat rail, the location clamping position of the needle will be changed. As a result, the uplink line smoothly penetrates an external bead provided by the uplink line bead transported device, as shown in Fig.17. Step 2: The bead threading device continues to move to the right. When the movable seat contacts the movable travel switch, the uplink line is just tightened. At this time, some of the motor operation will change as follows: Control motor A reversed means the bead threading device starts to move to the left. Controlling motor B rotated forward means the downlink line movable seat moves to the right for a suitable distance, providing two beads re quired for the next unit downlink line weaving. After that, controlling motor B stops. Linear motor F runs, the weaving beads are pushed into the weaving port and the output port, then moves back to the initial position. Linear motor E reversely drives, the pushing block of uplink line feeding bead returns to the initial position, and is on out feeding condition, as shown in Fig.18. S. Ouyang et al.2546 Step 3: Under the driving of the control motor A, the bead threading device moves to the left. When the threading device contacts the fixed stroke switch, some of the motor operation will change as follows: Linear motor G forward drives, the pushing block of the downlink line feeding bead pushes a shared bead to make the hole axis of the shared bead coincide with the needle axis. Control motor A rotated forward means the bead threading device starts to move to the right. Linear motor E drives forward, the push block of the uplink line feeding bead is in the feeding state. Control motor C rotated forward, the movable travel switch moves to the left for a suitable distance exactly equal to the length of the string required to weave every bead pad unit, as illustrated in Fig. 19. Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state." + ] + }, + { + "image_filename": "designv11_80_0002414_sii46433.2020.9025944-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002414_sii46433.2020.9025944-Figure6-1.png", + "caption": "Fig. 6: Definitions of joint angles w.r.t each joint", + "texts": [ + " Hence slopes of these links are equal to slope of the entire leg. Equating them gives the below equation zsj \u2212 zs(j+1) xsj \u2212 xs(j+1) = zs1 \u2212 zs5 xs1 \u2212 xs5 (22) where j = 1, 2, 3, 4 for each case Now since all the positions of the joints are known, the next step is to find out the joint angles which will formulate the Inverse Kinematics equations. Since the foot has to have complete contact with the ground, foot and the pelvis are parallel. Hence we get symmetry in the angles \u03b8s1 and \u03b8s5. The representation of all the angles is shown in the Fig. 6 . The exact equations of joint angles as given hereunder: \u03b8s1 = tan\u22121( xs1 \u2212 xs2 zs1 \u2212 zs2 ) (23) \u03b8s2 =| tan\u22121( ys3 \u2212 ys2 zs2 \u2212 zs3 ) | (24) \u03b8s4 =| tan\u22121( ys3 \u2212 ys4 zs3 \u2212 zs4 ) | (25) \u03b8s3 =| \u03b8s4 | + | \u03b8s2 | (26) \u03b8s5 = \u2212\u03b8s1 (27) These equations can generate the joint angles for any given start and destination positions. This method to solve inverse kinmatics is applicable if the position of CoM and the feet of the robot are known. The walking algorithm is only for given trajectory of CoM. This Inverse kinmatics solution can generalized for any small-sized humanoid robot having 5- DOF per leg with similar geometry" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001793_tmag.2019.2950181-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001793_tmag.2019.2950181-Figure3-1.png", + "caption": "Fig. 3. 3-D geometric model for (a) winding with a single conductor and (b) winding with triple sub-conductor.", + "texts": [ + " This voltage is added to the voltage equation in (5), yielding a final system of\u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 S + T d dt D D C d dt \u2212I RM C d dt MT 0 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1 \u23a3 a u i \u23a4 \u23a6 = \u23a1 \u23a3 0 0 U \u23a4 \u23a6 (24) where the additional matrix C is a horizontal assembly of matrices [C ,k]pn = \u2212lk 2\u03c0rk \u222b rad,k wnd (25) for all p and for k = naxi + 1, . . . , naxi + nrad, which account for the power coming from the radial slices. Note that C ,k = \u2212lk N DT ,k . (26) Equation (24) represents the whole MASM system, where the slices are coupled together. D. 3-D Simulation Time-harmonic 3-D FE simulation with COMSOL Multiphysics is used to validate the proposed MASM. Two test cases with npar = 1 and npar = 3 parallel conductors are simulated. The respective models are represented in Fig. 3. The multiple subconductor case provides better insight into frequency-dependent power losses in each conductor. The chosen geometries are symmetric with respect to the z = 0 plane, and thus, only the lower half z \u2264 0 is considered in the 3-D model for minimizing the required computation time and resources. Both halves are considered in the MASM. A boundary layer mesh is used in the conductor section for accurately capturing the influence of skin effect on power losses. Periodic boundary conditions are used on the sides of the symmetry sector" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000746_s40722-019-00139-6-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000746_s40722-019-00139-6-Figure1-1.png", + "caption": "Fig. 1 3D model of the WPC", + "texts": [ + ", \ud835\udf09 represent the displacement, velocity, and acceleration of the vessel; the indices \u201ck\u201d and \u201cj\u201d represent the j-mode oscillatory motion caused by the k-direction force; M is the rigid body mass matrix; A,B are the added mass and damping matrix, respectively; C is the restoring matrix; and F is the excitation force. For heave and pitch, the coupled 2-DOF equations are written as In this study, a 2.5D numerical method was used to predict the vertical motion performance of a high-speed catamaran. The parameters of the target ship are shown in Table\u00a01 and Fig.\u00a01 shows a body plan of the vessel. The hydrodynamic coefficients in the above equations are mainly related to three factors: the static buoyancy effect of the ship, the potential flow motion by the disturbance with free water surface, and the effect of the viscous flow on the naked hull. The hydrodynamic coefficients that are related to the static buoyancy of ships can be easily obtained, while those caused by the potential flow are mainly calculated by the panel method (Faltinsen et\u00a0al. 1991; Salvesen et\u00a0al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003423_iccsse50399.2020.9171958-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003423_iccsse50399.2020.9171958-Figure1-1.png", + "caption": "Figure 1. Containment vessel and polar crane: (1) Top head (2) End beam of bridge (3) Horizontal guide device (4) Girder of bridge (5) Trolley (6) Track beam (7) Crane traveling mechanism (8) Lifting equipment (9) Containment vessel.", + "texts": [ + " In view of the above problem, this paper proposed a synchronous control method based on the combination of synchronous relative coupling control and a novel nonsingular fast terminal sliding mode control to realize the convergence of velocity coupling error for a limited time, thus improving the positioning accuracy and antiinterference capability of polar crane. Finally, effectiveness and correctness of the control method are verified though MATLAB simulation. II. DYNAMIC ANALYSIS OF POLAR CRANE The polar crane is installed on the annular track above the containment vessel of the nuclear island, mainly including lifting mechanism, bridge frame, crane traveling mechanism, trolley and horizontal guide device, as shown in fig. 1. 978-1-7281-9846-0/20/$31.00 \u00a92020 IEEE Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 21,2020 at 06:45:56 UTC from IEEE Xplore. Restrictions apply. The crane traveling mechanisms drives the bridge frame to rotate, and the trolley makes linear motion on the bridge frame at the same time. Through the cooperative operation of the crane traveling mechanism and the trolley, the lifting equipment covers all areas in the containment vessel. Based on the principle of virtual power, a multi-rigid body dynamic model of polar crane is established" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002438_012018-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002438_012018-Figure1-1.png", + "caption": "Figure 1. The inertial and body frames of the quadcopter", + "texts": [ + " The rest of paper is organized as follows: Section 2 describes the mathematical model of the quadcopter, Section 3 gives a detailed description of proposed control method, Section 4 describes altitude and attitude controls by modified PID controller, Section 5 shows how to estimate the actual states by UKF, Section 6 gives a description of simulation block diagrams, Section 7 gives the results and discussion, and the conclusion is given in section 8. In this section, the mathematical dynamic model of the quadcopter is derived and can be used to study the performance of the quadcopter by using suitable simulation programs. A quadcopter has four independently controllable actuators (motors) assembled in cross configuration, as shown in Figure 1. One pair of opposite motors rotate in the same direction (clockwise), while the other two motors rotate counter clockwise, so the affected rotational torque on the body around Z axis of the quadcopter is cancelled during hovering state [8]. Figure 1 shows two basic coordinate frames, called an inertial frame (fixed earth frame(E)) and a body frame (B). These two frames are used to identify the quadcopter\u2019s location and attitude, and then The Fourth Postgraduate Engineering Conference IOP Conf. Series: Materials Science and Engineering 745 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/745/1/012018 translation and rotation matrices can be applied in order to transfer data from one coordinate frame to another [1, 8-10]. The quadcopter has six degrees of freedom in terms of linear position \u03be = (x, y, z) and the attitude (angular position) which is defined by Euler angles H = (roll (\u03d5), pitch (\u03f4), yaw (\u03c8)) with respect to the inertial frame" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001763_aeat-04-2019-0087-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001763_aeat-04-2019-0087-Figure5-1.png", + "caption": "Figure 5 Edge controller printed circuit board design 3D view", + "texts": [ + " Edge controller prototype development The proposed novel design of the edge controller avionics device consists of a multilayered PCB that may be enclosed in a protective cover. The current prototypes shown here do not use enclosures and expose the PCBs to both, the meteorological conditions of the flyingmachine and to observation from people examining the prototype. If an enclosure is to be engaged, proper ventilation has to be maintained for the cooling of the MOSFET transistors and their adjacent PCB area. A three-dimensional (3D) computer model of the edge controller PCB is presented in Figure 5. The dimensions of the PCB measure 65 30mm. The overall weight of a prototype board with all parts soldered is 10g. One can easily identify the six MOSFETs used in motor control \u2013 the 3 by 2 array of black rectangles in the closer-tothe-visitor PCB area. The large yellow blocks are tantalum capacitors that filter the power supply. Tantalum capacitors were chosen for their smaller dimensions and consequent lower air drag, compared to aluminum electrolytic polar capacitors. Further, the PCB hosts the MOSFET switching integrated circuits (ICs) \u2013 two eight-pin ICs were implemented" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002950_042044-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002950_042044-Figure2-1.png", + "caption": "Fig 2. The finite element simulation.", + "texts": [ + " Setting the appropriate solution step and time, the convergence range and other parameters[4], the torque calculation equation of ANSYS Maxwell software is: Vr V T T dd 2 1 2 1 = BJT (2.4) The torque of each unit nodes can be obtained from equation 2.4. The integration is the torque of the whole permanent magnet coupling can be obtained from post-processing the relationship curve between different parameters, the magnetic induction intensity of each region, and the eddy current density distribution, etc. Established the finite element analysis model from the physical drawing of the permanent magnet coupler as shown in Fig.2. The orange part is the conductor barrel, the red and blue parts are the permanent magnet blocks. The magnetization direction of the red permanent magnet block is opposite to the blue permanent magnet block are both along the radius direction[5]. In this paper, n42sh neodymium iron boron permanent magnet is used, which is composed of Nd2Fe14B. The maximum magnetic energy product is 398kj / m3, the maximum remanence BR is 1.47t, and the maximum coercive force is 1000ka / m. The Curie temperature of neodymium iron boron permanent magnet is 310 ~ 410 \u2103, and due to the poorly stability of temperature, the remanence temperature coefficient is about - (0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure82.9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure82.9-1.png", + "caption": "Fig. 82.9 CA GCI and AlSiC profile of von Mises stress (Model 3)", + "texts": [], + "surrounding_texts": [ + "\u2022 Suggested models of brake discs are of a solid design which needs to be pressed on to the axle before pressing of the wheels. For replacement of these brake discs, wheels are necessary to be pressed out, which is not required if the models are of a split type. Hence, further experiments may be done with split-type brake disc models. \u2022 Carbon matrix composites may be used to reduce the sound barrier and higher resistance to temperatures. \u2022 Brake pad analysis with various materials suitable to the disc material may be carried out to optimize pad material. 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 981 Appendix Structural load calculations: In this case, a railway vehicle travelling at a speed of 160 kmph on a horizontal track stops due to application of emergency brake was considered. Time of travel before stopping, deceleration, weight of the vehicle, clamping force on the brake disc, brake pad area, coefficient of friction, etc., for calculating the loads are taken from the railway specification. Value of Mass of railway vehicle\u2014M = 64000 kg, No. of axles per vehicle = 4, Maximum load per axle = 16000 kg, no. of brake discs per axle = 2, Load on each wheel = 8000 kg, Start speed v0 = 44.4 m/s, Deceleration a = 1.2 m/s2, Braking time ta = 36 s, Effective radius of the brake disc rdisc = 0.247 m, Radius of the wheel rwheel = 0.458 m, Mean coefficient of friction brake pad \u00b5 = 0.35, Clamping force Fc = 42.1 kN, Surface area of brake pads Ac = 400 cm2, Maximum temperature under sun = 70 \u00b0C, Maximum temperature under shade = 45 \u00b0C, Factor of Safety = 1.5 (Fig. 82.10). Stopping distance S \u00bc vots 1 2 at2s \u00bc 822:24m \u00f082:1\u00de Determination of pressure on disc Pressure acting on the brake disc, P \u00bc Fc Ac l \u00bc 42:1 1000 800 \u00f010\u00de 4 0:35 \u00bc 1:5Mpa on each side \u00f082:2\u00de 982 E. Madhusudhan Raju et al. where Fc Clamping force (i.e. 42 kN) Ac Contact area of brake pad on each side (i.e. 400 cm2) l Coefficient of friction (i.e. 0.35). Angular velocity x \u00bc Velocity radius \u00bc v0 rwheel \u00bc 44:44 0:458 \u00bc 97:12 rad=s \u00f082:3\u00de Thermal load The kinetic energy for one wheel (disc brake) is equivalent to the energy balance 0:125 1 2 M v2 \u00bc Zts 0 P\u00f0t\u00dedt \u00bc 2 Fdisc Zts 0 vdisc\u00f0t\u00dedt \u00f082:4\u00de The energy change at the moment is equal to the heat flux on the surface of the disc. Equation (82.4) is valid in the case of constant braking deceleration. The braking force on the disc is equal to Eq. (82.7) Fdisc \u00bc 0:125 1 2 M v20 2 rdisc rwheel v0 ts 1 2 a t2s \u00bc 8940N \u00f082:5\u00de The heat flux at the moment, which affects one half of the disc, is calculated according to the Q\u00f0t\u00de \u00bc Fdisc vdisc\u00f0t\u00de \u00bc Fdisc rdisc rwheel v0 a t\u00f0 \u00de \u00bc 8940 0:247 0:458 44:44 1:2 t\u00f0 \u00de \u00bc 214261 5786 tWatts \u00f082:6\u00de Area of friction surface \u00bc p 4 0:642 0:352 2 \u00bc 0:45m2 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 983 Q\u00f0t\u00de \u00bc 214261 5786 t 0:45 W/m2 \u00f082:7\u00de For the case of emergency braking on horizontal track from 160 kmph to stop, the analysis was carried out in 36 steps, each step being 1 s long." + ] + }, + { + "image_filename": "designv11_80_0000671_012029-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000671_012029-Figure1-1.png", + "caption": "Figure 1. Geometric parameters for a chain-wheel ensemble (a) [8], with details of first four contacts (b).", + "texts": [ + " In the case study, knowing the chain pitch (p), the diameter of the chain bush (db) and the number of teeth (z), the mathematical model allows the parameters of the bushed chain transmission and the numerical values of the contact point coordinates (xci, yci ) and also the contact angle (\u03b1i). 3. Determining the contact point and contact angle 3.1. Method I This first method consists in determining the contact angle (\u03b1i) for a number of contacts (i) between the chain bush and the sprocket. Based on this, the contact point coordinates (xi and yi) are calculated for the same number of contacts. The mathematical model for calculating the contact angle is based on the mounting position (see Figure 1), where tangential forces are assumed to be zero, and also that, for i = 0 result the first contact between the chain bush, and, the wheel position is vertical, where the notations considered to be with index 0 and \u03b1i = 0\u00b0. For a general case of chain-wheel ensemble, is considering i = 0, 1, 2\u2026, respectively j = i+1, the relation for determining the curves center position for sprocket (Ai, Aj) and also for chain-bush center PRASIC IOP Conf. Series: Materials Science and Engineering 514 (2019) 012029 IOP Publishing doi:10", + "1088/1757-899X/514/1/012029 2222 AiOAiOAiCiAiCi AiOAiCiAiOAiCi i yyxxyyxx yyyyxxxx cos . (28) 4. Case Study For considered chain teeth number (z), bush diameter (db) and chain pitch (p), the contact points coordinates between the chain bush and sprocket, which from a short-bush chain transmission, will be determined. So, for a standardized chain transmission with unitary ratio, are considered the teeth number z=16, the bush diameter db=5.08 [mm] and also the chain pitch p= 9.525 [mm]. The transmission parameters (Figure 1a) result as follow [7]: 8248 180 . z sin p Dd mm 7443.dDD bdf mm 682 3 069050501 . d .d.R b bmax mm 5524 2 1 .R D R max f A mm 6821 .Rr maxA mm 542 2 . d r b B mm. With the previous parameters and the first method presented, there are determined the contact angle (\u03b1) depending by the contact number (i) between the chain bush and sprocket. For contact angle known values (Figure 2), the contact points coordinates (xci, yci) can be determined, when the sprocket is rotating fron 0 to the angular pitch ( ,0 )", + " Analyzing the Figures 4, 5 and 6 can be observed that: the contact angle increases linear with the increasing of the sprocket rotation angle; the contact point coordinates on x axis decreases with the increasing of the sprocket rotation angle; the accentuate decrease is at contacts considered in increasing order; the contact point coordinates on y axis increases with the increasing of the sprocket rotation angle; the accentuate increase is at contacts considered in increasing order. PRASIC IOP Conf. Series: Materials Science and Engineering 514 (2019) 012029 IOP Publishing doi:10.1088/1757-899X/514/1/012029 For results comparison, in Table 6 are presented the point coordinates and also contact angles values. It is considered that, the first contact starts from initial position on the vertical axis, \u03b10 = 0, (Figure 1) and vary along the sprocket angular pitch, \u03c4 (Figure 3). PRASIC IOP Conf. Series: Materials Science and Engineering 514 (2019) 012029 IOP Publishing doi:10.1088/1757-899X/514/1/012029 The correct functioning of a bush chain transmission is given, among other functional parameters, by the contact point / angle between the bushing and the sprocket, knowing that with the increase of the contact angle the friction surface between them increases. The design of a chain drive with bushings must also take into account the contact angle or contact point between the bushing and the sprocket as they influence the transmission dynamics (vibrations and wear) and can compensate for deviations induced by assembly or manufacturing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003328_012034-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003328_012034-Figure2-1.png", + "caption": "Figure 2. Power screw with (a) square thread and (b) Acme thread.", + "texts": [ + " 6th International Week of Science, Technology and Innovation (6th IWSTI) Journal of Physics: Conference Series 1587 (2020) 012034 IOP Publishing doi:10.1088/1742-6596/1587/1/012034 Power screws are mechanical devices that change a rotation or angular displacement in a straight line, transmitting force and mechanical power. In practice, power screws are provided by specialized supplies that offer technical literature that includes all the data necessary for their selection. Two common types of thread are shown in Figure 2. Screwed joints can only transmit limited alternating stresses due to the notch stresses resulting from the functional design of a screw [12]. The square chord shown in Figure 2(a) provides the highest efficiencies and strengths; it also eliminates the radial force components between the bolt and nut. However, Acme chord is a common selection for power screws that must carry loads in both directions. The Acme cord, shown in Figure 2(b), has an included angle of 29\u00b0, which makes it easier to fabricate and also allows the use of a split nut that is tightened radially against the bolt to reduce wear. Most bolts are manufactured with only one cord (1 boot) [13]. 2.2.1. Coefficient of friction. The coefficient of friction must be greater than the tangent of the feed angle. For a static coefficient of friction on the screw flank of 0.15, the maximum corresponding feed angle should be 8.8\u00b0. In this case the coefficient of friction was maintained, and the feed angle was calculated [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003446_jae-200020-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003446_jae-200020-Figure1-1.png", + "caption": "Fig. 1. The PMSM prototypes, (a) canned, (b) ordinary.", + "texts": [ + " In this paper, a compensation element assisted thermal model is proposed and stressed, the compensation element works as amendment for deviation caused by heat flow, with emphasizing the influence of slot effect and distributed winding, as well as formula inference and the establishment of thermal network model. The 3D FE is used to validate the novel thermal model. Recently, permanent magnet synchronous motor (PMSM) is an attractive drive candidate\u00a0[17\u201319], and some canned PMSM motors are also reported in pump industry\u00a0[20,21]. In Fig.\u00a01 a couple of motors, with identical geometry except one of them is canned, are shown. Figure\u00a01a shows the canned motor with distributed winding, this winding topology shows characteristics that are as follows: (1)\u00a0more sinusoidal airgap flux field; (2)\u00a0lower harmonic electromagnetic force. Consequently, this topology is suitable for the canned motor. Compared with Fig.\u00a01b, a couple of cans are used in the airgap, namely the stator can and rotor can. The stator can is fixed into the inner surface of the stator teeth to protect the armature windings, and the rotor can is fixed into the outer surface of the permanent magnets to protect magnets from corrosion, meanwhile to reduce hydraulic friction. Geometrical specifications are listed in Table\u00a01. To accommodate PMs and cans, the airgap width with PMs enlarges to 7.9\u00a0mm. Because airgap width is artificially increased to install cans, the mutual induction can be neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure5.11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure5.11-1.png", + "caption": "Figure 5.11 Scheme of a tunnel cross-section and designations of stresses in different coordinate systems.", + "texts": [ + " The stresses which act on mutually-perpendicular areas n = const. and t = const. can be expressed through the stresses, acting on the areas r = const., \ud835\udf03 = const.: \ud835\udf0e\ud835\udc5b\ud835\udc5b = \ud835\udf0e\ud835\udc5f\ud835\udc5fcos2\ud835\udefc + \ud835\udf0e\ud835\udf03\ud835\udf03sin2\ud835\udefc + 2\ud835\udf0f\ud835\udc5f\ud835\udf03 sin\ud835\udefc cos\ud835\udefc; \ud835\udf0e\ud835\udc61\ud835\udc61 = \ud835\udf0e\ud835\udc5f\ud835\udc5fsin2\ud835\udefc + \ud835\udf0e\ud835\udf03\ud835\udf03cos2\ud835\udefc \u2212 2\ud835\udf0f\ud835\udc5f\ud835\udf03 sin\ud835\udefc cos\ud835\udefc; (5.5.2) \ud835\udf0f\ud835\udc5b\ud835\udc61 = (\ud835\udf0e\ud835\udf03\ud835\udf03\u2212\ud835\udf0e\ud835\udc5f\ud835\udc5f) sin\ud835\udefc cos\ud835\udefc + \ud835\udf0f\ud835\udc5f\ud835\udf03(cos2\ud835\udefc \u2212 sin2\ud835\udefc). The corresponding relationships for the displacements take the form: U\ud835\udc5b\ud835\udc5b = Ur cos\ud835\udefc \u2212 U\ud835\udf03\ud835\udc5f sin\ud835\udefc; (5.5.3) U\ud835\udc61\ud835\udc61 = Ur sin\ud835\udefc + U\ud835\udf03 cos\ud835\udefc, where \ud835\udefc \u2013 the angle between the straight line \ud835\udf03 = const., presenting the direction of a normal at the given point in Figure 5.11. The angle \ud835\udefc can be determined by the formulas: tan\ud835\udefc = k2 \u2212 k1 1 + k1k2 ; cos\ud835\udefc = 1 + k1k2\u221a 1 + k2 1 \u221a 1 + k2 2 ; sin\ud835\udefc = k2 \u2212 k1\u221a 1 + k2 1 \u221a 1 + k2 2 (5.5.4) k1, k2 \u2013 angular coefficients of the corresponding straight lines Figure 5.11. If the contour of the underground structure is described by the function y = f(x), then the angular coefficients, which correspond to the point on the curve with the coordinates xi and yi, can be determined as follows: k1 = y(xi) xi ; k2 = \u2212 1 y\u2032(xi) (5.5.5) 5.5 Calculations of Underground Structures 309 The boundary conditions are formulated as follows. If the given problem is about an impact of a seismic wave on a non-reinforced tunnel of an arbitrary cross-section, then the satisfaction of the following conditions is necessary: \ud835\udf0e\ud835\udc5b\ud835\udc5b(ri, \ud835\udf03i) = 0; \ud835\udf0f\ud835\udc5b\ud835\udc61(ri, \ud835\udf03i) = 0; i = 1, 2, \u2026K (5", + " Since the assigned coefficients determined by expressions (5.5.1) are complex quantities, the unknown coefficients An and Bn are also complex. Therefore, to solve the equations we equate the real and complex expressions, and instead of two equations of the form (5.5.7) we obtain four such equations. For evaluating the accuracy of the proposed method the calculations of an unsupported tunnel with elliptical cross-section under the action (influence) of an elastic longitudinal wave, propagating in the Ox and Oy directions were conducted (Figure 5.11). The reason for picking such method of calculations is due to the fact that for this case numerical results were obtained by another method \u2013 the perturbation forms of boundary method [195]. The equation of the ellipse in the Cartesian coordinates can be written in the form of: y = b \u221a 1 \u2212 (x\u2215a)2 (5.5.8) where a, b \u2013 semi-axis (half axis) of ellipsis. The angle \ud835\udefc is determined from the expression: cos\ud835\udefc = (x a )\u22122 {1 + (b a) 2 [1 \u2212 (x a )2 ] (x a )\u22122 } \u22121\u22152 \u00d7 {1 + (x a )\u22122 (b a) \u22122 [1 \u2212 (x a )2 ]} \u22121\u22152 (5" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000996_ivs.2019.8813855-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000996_ivs.2019.8813855-Figure3-1.png", + "caption": "Fig. 3. Variable definitions, the indices stand for: tire, bike, human, f ront and rear", + "texts": [ + " A full suspension mountain bike with electronically adjustable dampers and an ABS system (Fig. 2) has been equipped with an extensive set of sensors to perform an in-depth analysis of the relevant effects and to validate the simulations. The sensors evaluated in this paper include: a pressure sensor mounted on the hydraulic front disk brake, incremental wheel speed sensors and inertial measurement units (IMU) measuring acceleration and angle rates in three axes. The sensor locations and a definition of the variables and coordinates can be found in Fig. 3. Longitudinal and vertical deflection is denoted by x and z, \u03c9 describes angle rates, a acceleration and \u03d5b the fork bending angle. The simplest conceivable model of vertical vehicle dynamics is the 1DOF (one degree of freedom) quarter-car model shown in Fig. 4(a). The major assumption is that the tire perfectly follows the road profile (zr = zt), resulting in a simple second order transfer function: G(s) = zb zr = c/k s+ 1 M/k s2 + c/k s+ 1 (1) The system can be characterized by its eigenfrequency \u03c90 = \u221a k/M and damping ratio D = c/(2 \u221a Mk)", + " 4(d): x\u0307 = Ax + Bu x = ( zh zb z\u0307h ( z\u0307b \u2212 c Mb zt ))> , u = zr = zb A = 0 0 1 0 0 0 0 1 kh Mh \u2212 kh Mh ch Mh \u2212 ch Mh \u2212kh+k Mb \u2212 kh Mb \u2212 ch+c Mb ch Mb B = ( c Mb 0 chc Mh/Mb ( k Mb \u2212 c(ch+c) M2 b ))> (2) In [23] the rider suspension properties have been analyzed by test bench measurements and modeled by a multibody model. The rider was seated on a rigid bicycle saddle, holding a handlebar that was vertically excited by vibrations. The rider response was measured optically in the zh-direction, as visualized in Fig. 3. For the quarter-car approach in this section the recorded experimental data is fitted by a simple spring-mass-damper system as seen in Fig. 5, yielding a rider eigenfrequency of 3 Hz and a damping value of 0.5. Note that the coordinate axes of input and output were not aligned. Therefore, the experimental data has been scaled in order to achieve unity gain at 0 Hz, which corresponds to a purely vertical spring-damper model. The 1DOF and the 2DOFh models allow an input output simulation as in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002970_012003-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002970_012003-Figure1-1.png", + "caption": "Figure 1. Electrospark alloying process scheme: (1) alloy electrode; (2) hardenable product; (3) platten; (4) application head; (5) servomotor of tracking system; (6) gearbox; (7) feed-screw.", + "texts": [ + "), metals (Ti, Cr, Co, Ni, Al, W, Mo, Re, Ta, Hf , etc.), as well as alloys based on them. The required value of the interelectrode gap is supported by the movement of the electromechanical tracking system (U). In order to increase the flatness of the hardened layer and uniform erosion of the alloying electrode, it is given rotation from an electric motor with a frequency of 400 ... 4000 rpm. The alloying electrode with a diameter of 0.5 ... 2.0 mm is installed in a collet chuck. The alloying process scheme is shown in figure 1. Alloy electrode (1) is an anode, and hardened product (2), installed on the desktop (3), the cathode. The parameters of electric spark discharges are selected so that predominant erosion of the alloying electrode occurs. The necessary power, frequency and duration of the discharges are provided by a pulse generator, which supplies rectangular voltage pulses to the doping electrode. The required value of the interelectrode gap is supported by an electromechanical servo system, by moving the servo head (4) in the vertical direction by means of an actuator motor (5), a worm gearbox (6) and a lead screw (7)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000864_012083-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000864_012083-Figure1-1.png", + "caption": "Figure 1. Permissible forms of contact spots a) permissible location of the contact patch, b) types of permitted forms of contact spots.", + "texts": [], + "surrounding_texts": [ + "Content from this work may be used under the terms of the Creative Commons Attribution 3.0 licence. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI.\nPublished under licence by IOP Publishing Ltd\nISTC-IETEM\nIOP Conf. Series: Materials Science and Engineering 570 (2019) 012083\nIOP Publishing\ndoi:10.1088/1757-899X/570/1/012083\nBevel gears monitoring methods developmentin the total contact patch terms\nD T Safarov, A G Kondrashov\nNaberezhnye Chelny Institute of Kazan (Volga Region) Federal University, Naberezhnye Chelny, Mira Ave. 68/10, 423800, Russian Federation Abstract. The article describes the main provisions of the method, which allows to perform the control of the gear rims of bevel gears according to the shape and relative position of the total contact patch. In comparison with the standard, the developed method allows to evaluate the conformity of the teeth of the ring gear with its considerable displacements relative to the calculated point of the gear transmission.\n1. Introduction. Conical spur gears in the construction of heavy vehicles are used in the differential mechanism, allowing to redistribute the torque from the wheels, when cornering and in the conditions of wheel slip when the vehicle is moving on dirt roads.\nTo ensure the durability of the differential operation, it is important to ensure the required level of contact patch over all teeth of bevel gears included. Irregularity of the area and location of contact spots on the teeth of the rims of the conical wheels can lead to uneven loading of individual teeth and their intense local wear. In some most unfavorable operating conditions of the car, their breakdown is possible.\nThe difficulty of ensuring a given spot size of the contact bevel gear in the differential mechanism lies in the fact that it is unregulated. Many factors influence the contact patch size, beginning with the design stage, for example, cutting tool profile provision [1,2], and at the stage of cutting tools wear [3], processed layer state of the blank [4] and other factors. As a result of the above factors, the teeth of the ring gear receive different geometrical deviations, leading to deviations of the location and area of the gear contact patch. Received deviations of the location and area of contact spots can\u2019t be compensated or redistributed by means of adjusting the position of gears in the process of assembling the differential. Consequently, the operations of bevel gears are imposed special requirements of ensuring the total contact patch on (picture1).\nUnder workshop conditions, the main method for assessing the size of the contact patch is a visual method of comparing the resulting contact patch with the allowed forms. Measurements and recording of its parameters are not provided. A visual comparison of the acceptable forms and contact spots location by experienced and qualified personnel ensures the availability of product, but it does not allow to register the quantitative values of this indicator, to monitor the trends of its changes, and therefore effectively manage the process of manufacturing wheels.", + "ISTC-IETEM\nIOP Conf. Series: Materials Science and Engineering 570 (2019) 012083\nIOP Publishing\ndoi:10.1088/1757-899X/570/1/012083\n2. Theoretical part The contact spot requirements for bevel gears are established by a number of regulatory documents [5, 6]. Regulatory requirements are formulated in tolerance form (Figure 2 a). Standard requirements [1] are applied to the total contact patch dimensions. Its size is determined as a percentage of the length of the tooth \u201cthe ratio of the distance between the extreme points of traces of fit to the length of the tooth\u201d (\n\u00d7 100), the height of the tooth \"the ratio of the average height of the traces of fit to the average height\nof the tooth of the corresponding active lateral surface\" ( \u00d7 100) (Figure 2 b).\nDetermining the size of the total contact patch implies its initial location in the coordinates of finding the calculated point (at the intersection of the average height of the tooth and forming the separating cone), i.e. in the conditions of already adjusted process of a crownwork of a cogwheel crown.\nIn practice, even in a well-established process of teeth processing, the total contact patch can significantly shift relative to the calculated gearing point (Figure 2c), and in some cases, may go beyond the boundaries of the teeth. Under these conditions, direct adherence to the method of calculating and rationing the dimensions of the total contact patch by the standard method is the risk of declaring unfit products suitable. Thus, there is a need to develop a more universal method for estimating the total contact patch.\nThe developed method is given additional modern requirements concerning its incorporation into the quality system of the enterprise [7], providing a collection of quantitative data on controlled characteristics, the identification of diagnostic components of the process of tooth processing [8], suitable for evaluating the result of processing as a result of mathematical modeling of the process [9]. As a result,", + "ISTC-IETEM\nIOP Conf. Series: Materials Science and Engineering 570 (2019) 012083\nIOP Publishing\ndoi:10.1088/1757-899X/570/1/012083\nto ensure the above requirements, in the developed method, the form of the total contact patch and its relative positioning are separately normalized. To establish the requirements of the contact patch shape, similar to the standard method, t the minimum and maximum limits of its boundary are set (Figure 3 b). In contrast with it, the nominal or ideal position of the total contact patch is given as an ellipse. For the considered bevel gear, according to the test results, its optimal position was found, shifted from the calculated transfer point towards the flat end (Figure 3 a). The minimum and maximum dimensions of the ellipse are set depending on the length of the tooth along the pitch circle and the average height of the tooth.\nThus, the assessment of compliance of the total contact patch is simultaneously performed on a group of indicators: % L,% H - normalizing the shape of the total contact patch along the tooth and its height, as well as indicators L1, L2, L3, L4 - setting restrictions on the relative location of the contact spot. For each of these indicators, the upper and lower deviations are calculated.\nFinding the actual values of these indicators is carried out according to the graphical processing of photographic images of contact patches formed as a result of erasing ferric acid pigment on the lateral surfaces of the ring gear teeth (ls) as a result of rolling the processed gear and the reference gear (Figure 2c)." + ] + }, + { + "image_filename": "designv11_80_0002935_0954408920932358-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002935_0954408920932358-Figure3-1.png", + "caption": "Figure 3. Single pipeline model and forces diagram of magnet.", + "texts": [ + " Solve the equations of induced electromotive force The design of energy collector is based on Faraday\u2019s electromagnetic induction law. When the magnet passes through the coil, the energy will be generated in the coil. In fact, the study of all the power generation devices can be illustrated by analyzing the single one element pipeline, because the energy collecting device is symmetrical in structure and it has the same principle as the two element pipeline model. Schematic diagram of single pipe model is shown in the Figure 3. From Figure 3, we can see the working situation of the device which include gravity (mg), friction force (Ff), support force (FN), elastic force (FT), electromagnetic drag force (Fem). We can infer the equations of magnet from the Newton\u2019s law: mat \u00bc mx00m \u00bc mg cos h Fem Ff \u00fe FT FN Fc \u00bc mx2r (1) The total force of magnet offsets the centrifugal force, which makes the magnet has constant velocity. According to the Faraday\u2019s law of electromagnetic induction, the equation of induced voltage can be given by21: e \u00bc a\u00f0r\u00de r0 (2) From equation (2), a\u00f0r\u00de, named coupling factor, is relative to magnetic field intensity (B) and the length of closed coil (lw)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003742_icarm49381.2020.9195386-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003742_icarm49381.2020.9195386-Figure4-1.png", + "caption": "Fig. 4. Simplified kinematic model.", + "texts": [ + " Since the motion of ankle and foot complex is mainly affected by chain 1, in the following motion analysis, chain 1 is focused on. Actually, the objective of motion analysis is to calculate velocity transmission ratio between Link 1 and Link 9. During normal walking, the rotation motion mainly occurs at the ankle joint. Therefore, in this analysis, in order to simplify the derivation, the motion of subtalar joint is neglected. Based on above simplifications, the kinematic diagram of the motion system can be described as Fig. 4. The system has four links, Link 1\u2019, 2\u2019, 3\u2019 and 4\u2019; they are equivalent to Base, Link 1, Link 2 and Link 9 in Fig. 3 respectively. Additionally, the system consists of two rotational joints and two spherical joints; they are R1, R2, S1 and S2 respecitvely. Accordingly, four coordinate systems are defined. Specifically, as shown in Fig. 4, o1, o2, o3 and o4 are the original points of the 978-1-7281-6479-3/20/$31.00 \u00a92020 IEEE 583 Authorized licensed use limited to: University of New South Wales. Downloaded on November 15,2020 at 09:08:52 UTC from IEEE Xplore. Restrictions apply. four coordinate systems respectively. Based on method of Denavit-Hartenberg (DH) presentation [15], coordinate axis k1 and k4 are set along the axial directions of revolute joint R1 and R2 respectively. k2 is parallel to k1 and passes the center of spherical joint S1; while k3 passes through the centers of spherical joint S1 and S2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001535_ecce.2019.8912634-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001535_ecce.2019.8912634-Figure3-1.png", + "caption": "Fig. 3. Winding arrangement of target PMSM", + "texts": [ + " The reference map is not changed by the DC voltage. Fig. 2 shows modulation degree maps for 260 V and 380 V. The one for 260 V is 1.46 times as large as that for 380 V because of the DC voltage fluctuation and the same current reference. The main parameters of a PMSM used for testing are summarized in Table I. The test PMSM is an 8-pole 48-slot IPMSM with distributed winding. The structural features of the PMSM include a structure for reducing torque ripple focused on the winding factor of the PMSM and a magnetic gap [9]. Fig. 3 shows the winding arrangement of the target PMSM. The carrier frequency is 8.5 kHz (f1: 800 Hz) to output about 10 pulses per 1 electrical angle at 12,000 rev/min. The voltage reference update timing is 1/fc second (fc: Carrier frequency) to reduce the microcomputer load. The voltage reference update timing is the most important factor for determining the microcomputer performance. Radial electromagnetic forces are considered to be a major contributor to radiated magnetic noise and stator vibrations" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001881_s12206-019-1113-4-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001881_s12206-019-1113-4-Figure8-1.png", + "caption": "Fig. 8. Measurement positions for hardness test.", + "texts": [ + " Moreover, \u03c3t / \u03c3max remained high against the helix angle near ACUTE-END in MESH-A. For instance, \u03c3t / \u03c3max was approximately 0.95 for \u03b20 = 20.00\u00b0 and approximately 0.75 for \u03b20 = 40.88\u00b0 near ACUTE-END. Therefore, we expect that the hardened layer near ACUTE-END influences the bending fatigue strength in MESH-A. We measured the Vickers hardness at the center of the tooth width of a test gear for each helix angle. We cut off each target gear in the tooth normal direction at 10 mm from the gear side surface using a precision cutting machine, as shown in Fig. 8(a). We measured the hardness along Hofer\u2019s critical cross section on this cut surface as shown in Fig. 8(b). The measurement direction was the normal direction of the cut surface. We used a Vickers hardness tester (Mitutoyo HM102) for this hardness test. All the test gears used for this hardness test were the gears of CASE-TS type. Note that the type of case-carburizing of gear side surfaces (CASE-T or CASE-TS) has little influence on this hardness test because the measurement points are far from the tooth width ends. We measured the Vickers hardness at ACUTE-END of a test gear for each helix angle in CASE-T. The measurement points were located along Hofer\u2019s critical cross section at ACUTE- END as shown in Fig. 8(a). In the same way, we also measured the Vickers hardness at ACUTE-END for CASE-TS. We measured the residual stress on the critical cross sections of the teeth faces using an X-ray residual stress analyzer (PULSTEC \u03bc-X360s). As the residual stress in CASE-TS, we measured the residual stress of a supporting gear for each helix angle. We measured the residual stress of four teeth of each target gear. We measured the residual stress at 13 measurement points: 0.4 mm, 2.8 mm, 5.2 mm, 7.6 mm, 10.0 mm, 12" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000542_978-3-030-20131-9_261-Figure76-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000542_978-3-030-20131-9_261-Figure76-1.png", + "caption": "Fig. 76. Schematic diagram of string installation.", + "texts": [], + "surrounding_texts": [ + "In chapter 2, three methods of automatic knitting of bead mat are put forward. Among them, Warp and Weft Automatic Weaving Method is the most simple, but the cost is higher and its stability is poor. The lock stitch sewing weaving method is rea lly good, but there exists three difficulties applied in the machine: The cross-sectional size of the hook must be smaller than the size of the bead hole as the hook needle needs to completely pass through the bead, which makes it difficult to hook the string smoothly. It is difficult to guarantee the parallelism between the cross-section of the loop ring and the end of the bead of the first row in steps 4-6. When weaving a larger size beaded cooling pad, threading the string into the longer aligned transverse bead holes becomes hard. So compared with the other two methods, single-line straight-through method, which has good stability and high knitting efficiency and is easy to be realized on the machine, is the focus of the discussion below. Based on this method, an automatic weaving device capable of weaving a beaded cool pad is proposed and designed. 4.1 Feeding Device Design Before weaving the beaded cooling pad, since all the beading arrangements are disordered, it is necessary to design a device that puts the beads into the weaving state in an orderly manner, which is called feeding device. Referring to the feeding mechanism of the firecracker weaving [14-16], Fig.14 shows the schematic diagram of the designed bead feeding mechanism device. Before the device runs, all the beads are placed in the hopper and the two guiding wheels. A small number of longitudinally are placed in a horizontal arrangement. When the two guide wheels rotate in opposite directions, the beads in the hopper will be putted into the guide groove and conveyed to the front of the beading device in an orderly manner. In order to avoid a rigid collision between the feeding device and the ball transported device, the end of the guiding groove is made by a material with better elasticity. 4.2 Design and Working Principle of the Beaded Pad Weaving Device Fig. 15 demonstrated the beaded pad weaving device. Since the figure is only for explaining the movement process of the beaded pad weaving device, the feeding device is not shown in the figure. And there are 7 motors in this device. Control motor A controls the movement of the threading device. Control motor B controls the movement of the movable line of the downlink line. Control motor C controls the movement of the movable stroke switch. Control motor D controls the rotation of the output port and the braided port. Linear motor E S. Ouyang et al.2544 drives the up-line feed ball push block movement. Linear motor F pushes the braided beaded cool pad unit into the braided port. Linear motor G drives the linear motion of the downlink line feed bead block. Before the device is operated, the downlink threading is first performed. After the downlink threading is finished, the string is installed into the beaded pad weaving device. The downlink line with heavy beads at the end is wrapped around the fixed pulley mounted on the frame. Then it passes through the downlink line movable seat, the through hole, the downlink end sleeve, the bead and the braided port successively, to reach the uplink line. And the uplink line is directly connected to the needle of the threading device, the first string of beads are moved to the corresponding position on the weaving port, as shown in Fig.16. After the string installation is completed, the motor A is manually controlled to make the driving roller be located between the two trapezoidal blocks on the movable seat rail of the threading device. The manually controlled linear motor E is to drive the uplink line to send th e beads push block moves, which is external bead conveyed from the feeding device. It causes the holes axis of the bead to coincide with the needle axis of the bead threading device and the up-line bead push block is in the beading state. The specific work ing process of the device is as follows: Method Research and Mechanism Design of Automatic Weaving\u2026 2545 Step 1: Under the drive of the control motor A, the bead threading device moves to the right. When the driving roller moves in the second trapezoidal block on the movable seat rail, the location clamping position of the needle will be changed. As a result, the uplink line smoothly penetrates an external bead provided by the uplink line bead transported device, as shown in Fig.17. Step 2: The bead threading device continues to move to the right. When the movable seat contacts the movable travel switch, the uplink line is just tightened. At this time, some of the motor operation will change as follows: Control motor A reversed means the bead threading device starts to move to the left. Controlling motor B rotated forward means the downlink line movable seat moves to the right for a suitable distance, providing two beads re quired for the next unit downlink line weaving. After that, controlling motor B stops. Linear motor F runs, the weaving beads are pushed into the weaving port and the output port, then moves back to the initial position. Linear motor E reversely drives, the pushing block of uplink line feeding bead returns to the initial position, and is on out feeding condition, as shown in Fig.18. S. Ouyang et al.2546 Step 3: Under the driving of the control motor A, the bead threading device moves to the left. When the threading device contacts the fixed stroke switch, some of the motor operation will change as follows: Linear motor G forward drives, the pushing block of the downlink line feeding bead pushes a shared bead to make the hole axis of the shared bead coincide with the needle axis. Control motor A rotated forward means the bead threading device starts to move to the right. Linear motor E drives forward, the push block of the uplink line feeding bead is in the feeding state. Control motor C rotated forward, the movable travel switch moves to the left for a suitable distance exactly equal to the length of the string required to weave every bead pad unit, as illustrated in Fig. 19. Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state." + ] + }, + { + "image_filename": "designv11_80_0003377_ccdc49329.2020.9164095-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003377_ccdc49329.2020.9164095-Figure2-1.png", + "caption": "Figure 2: The Shape Parameter of SAW in DATCOM", + "texts": [], + "surrounding_texts": [ + "d\u03072k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)k(ld sin \u03b4k)\u03b4\u0307k \u2212(l cos \u03b4k)\u03b4\u0307k\n\u23a4 \u23a6 ,\nand\nd\u03082k\u22121(ld) =\n\u23a1 \u23a3\n0 0 0\n\u23a4 \u23a6 ,\nd\u03082k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)kld(cos \u03b4k \u03b4\u0307 2 k + sin \u03b4k \u03b4\u0308k)\nld(sin \u03b4k \u03b4\u0307 2 k \u2212 cos \u03b4k \u03b4\u0308k)\n\u23a4 \u23a6 .\nBecause the first derivative of the \u03b4k represents the folding angular rate, and in the assumptions presented above, the folding angular rate during morphing process keeps constant, so the second derivative of \u03b4k is zero, thus the second derivative of d2k can be expressed as follows:\nd\u03082k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)kld(cos \u03b4k)\u03b4\u0307 2 k\n(ld sin \u03b4k)\u03b4\u0307 2 k\n\u23a4 \u23a6 . (6)\nAfter computing d\u0307 and d\u0308 in equation (2) and (3) above, Meanwhile, the domain item dcm of Fext in (2) can also be decomposed in two parts:\nd\u0308cm = d\u0308cm1 + d\u0308cm2, (7)\nand\ndcm1 = 1\n2m mf lf (R1 +R3),\ndcm2 = 1\n2m mtipltip(R2 +R4),\n(8)\nwhere Ri(i = 1, 2, 3, 4) are represented as follows:\nR1 = 0, R2 = \u03b4\u030721 [0,\u2212 cos \u03b41, sin \u03b41] T , R3 = 0, R4 = \u03b4\u030722 [0, cos \u03b42, sin \u03b42] T ,\nSimilarly, the additional moment caused by morphing of wing tips can be computed. The domain item ( \u222b [d\u0304]d\u0308dm) in Mext can be given as follows: \u222b\n[d\u0304]d\u0308dm = 1\n2 [mf lf (S1 + S3)\n+ mtipltip(S2 + S4)][1, 0, 0] T , (9)\nwhere Si(i = 1, 2, 3, 4) represent the additional terms caused by folding of wing tip:\nS1 = 0, S2 = (l0 \u2212 ltip)\u03b4\u0307 2 1 sin \u03b41,\nS3 = 0, S4 = (ltip \u2212 l0)\u03b4\u0307 2 2 sin \u03b42,\nWhere l0 is the half width of the front view fuselage. Then, substituting the domain item d\u0308cm in Fext and\nthe domain item \u222b [d\u0304]d\u0308dm in Mext into the equation above, the Fext and Mext can be given as follows:\nFext = \u22121\n2 mtipltip(R\u03b41\u0394\u03b41 +R\u03b42\u0394\u03b42), (10)\nwhere R\u03b41 = \u03b4\u030721 [0, sin \u03b41, cos \u03b41] T ,\nR\u03b42 = \u03b4\u030722 [0,\u2212 sin \u03b42, cos \u03b42] T ,\nand\nMext = \u22121\n2 mtipltip(S\u03b41\u0394\u03b41 + S\u03b42\u0394\u03b42)[1, 0, 0]\nT , (11)\nwhere S\u03b41 = (l0 \u2212 ltip)\u03b4\u0307 2 1 cos \u03b41,\nS\u03b42 = (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42.\nThen, substituting equation (10) and (11) into equation (2) and (3), the nonlinear model of the folding wing-tip UAV can be expressed as follows:\nu\u0307 = rv + qw \u2212 g sin \u03b8 + Fx\nm , (12)\nv\u0307 = \u2212ur + wp+ gcos \u03b8sin\u03c6+ Fy\nm\n\u2212 mtipltip 2m (\u03b4\u030721 sin \u03b41\u0394\u03b41 \u2212 \u03b4\u030722 sin \u03b42\u0394\u03b42),\n(13)\nw\u0307 = uq \u2212 vp+ gcos \u03b8cos\u03c6+ Fz\nm\n\u2212 mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42),\n(14)\np\u0307 = (c0r + c1p)q + c2L\u0304+ c3N \u2212 c4 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42),\n(15)\nq\u0307 = c5pr \u2212 c6(p 2 \u2212 r2) + c7M, (16)\nr\u0307 = (c8p\u2212 c2r)q + c4L\u0304+ c9N, (17)\nwhere \u03b4\u03071,2 is the folding rate of the wing tip, \u0394\u03b41,2 is the folding angle change in a period of folding process, L\u0304,M ,N are components of total moment along three axes within body frame and can be given as follows:\nL\u0304 =p\u0307Ix \u2212 r\u0307Ixz + qr(Iz \u2212 Iy)\u2212 pqIxz \u2212 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42)\nM = Iy q\u0307 + pr(Ix \u2212 Iz) + (p2 \u2212 r2)Ixz\nN = r\u0307Iz \u2212 p\u0307Ixz + pq(Iy \u2212 Ix) + qrIxz\nwhere ci(i = 0, 1, 2, ..., 9) are constant coefficients expressed as follows:\nc0 = ( (Iy \u2212 Iz)Iz \u2212 I2xz\nIxIz \u2212 I2xz ), c1 = (Ix \u2212 Iy + Iz)Iz \u2212 Ixz IxIz \u2212 I2xz\nc2 = Iz\nIxIz \u2212 I2xz , c3 = Iz IxIz \u2212 I2xz , c4 = 1 IxIz \u2212 I2xz\nc5 = Iz \u2212 Ix\nIy , c6 = Ixz Iy , c7 = 1 Iy ,\nc8 = Ix(Ix \u2212 Iy) + I2xz\nIxIz \u2212 I2xz , c9 = Ix IxIz \u2212 I2xz\n1816 2020 Chinese Control And Decision Conference (CCDC 2020)\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply.", + "And Fx,Fy ,Fz are components of aerodynamic force and propulsive force along three axes within body frame, given that the proplusive force is along body x-axis and the engine offset angle \u03b1T = \u03b2T = 0, namely, T = Tx. Based on the conversion relationship between the body frame and airflow frame Xbody = ST \u03b1\u03b2Xwind, there exists:\n\u23a1 \u23a3 Fx\nFy Fz \u23a4 \u23a6 body = \u23a1 \u23a3 T 0 0 \u23a4 \u23a6 body + ST \u03b1\u03b2 \u23a1 \u23a3 \u2212D Y \u2212L \u23a4 \u23a6 wind (18)\nHence, Fx,Fy ,Fz are given as follows:\nFx=T + L sin\u03b1\u2212 Y cos\u03b1 sin\u03b2 \u2212D cos\u03b1 cos\u03b2, (19)\nFy = Y cos\u03b2 \u2212D sin\u03b2, (20)\nFz=\u2212L cos\u03b1\u2212 Y sin\u03b1 sin\u03b2 \u2212D sin\u03b1 cos\u03b2, (21)\nwhere L,Y ,D are thrust, side and drag force. \u03b1,\u03b2 are angleof-attack and the sideslip angle, \u03b8 and \u03c6 are pitch angle and roll angle.\nWith decoupling method proposed in [21], the decoupled longitudinal nonlinear model is given as follows:\nmV\u0307 = T cos\u03b1\u2212D+mg(\u2212cos\u03b1 sin \u03b8+sin\u03b1 cos \u03b8), (22)\nmV \u03bc\u0307 = T sin\u03b1+ L\u2212mg(sin\u03b1 sin \u03b8 + cos\u03b1 cos \u03b8)\n\u2212mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (23)\nmV \u03b1\u0307 = \u2212T sin\u03b1\u2212L+mV q+mg(sin\u03b1 sin \u03b8+cos\u03b1 cos \u03b8)\n\u2212mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (24)\n\u03b8\u0307 = q, (25)\nIy q\u0307 = M, (26)\nSimilarly, the decoupled lateral nonlinear model is given as follows:\nmV \u03b2\u0307 = Y \u2212mV (\u2212p sin\u03b1) + r cos\u03b1\n\u2212 mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (27)\n\u03c6\u0307 = p+ (r cos\u03c6+ q sin\u03c6) tan \u03b8, (28)\n\u03c8\u0307 = 1\ncos \u03b8 (r cos\u03c6+ q sin\u03c6), (29)\np\u0307 = (c0r + c1p)q + c2L\u0304+ c3N \u2212 c4 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42),\n(30)\nr\u0307 = (c8p\u2212 c2r)q + c4L\u0304+ c9N, (31)\nIt can be obtained from the equation (12)-(17) that during the morphing process of folding wing-tip UAV, some aerodynamic items such as the x-axis direction component of airspeed V , pitch and yaw angular rate, seem to have no distinction with that of conventional aircrafts. Consequently, It can be derived from above that the whole morphing process, including taking off, taking up and cruise, only has impacts on the aerodynamic performance of the folding wing-tip UAV in y-axis or z-axis, and in x-axis, the UAV appears the same as conventional aircrafts.\n3 Numerical Simulation\nTo validate the utility of the modeling method and nonlinear models of the folding wing-tip UAV, a numerical study is conducted to SAW, which is a typical kind of folding wing-tip UAV. The 3D model of the SAW is established in DATCOM and aerodynamic performances under different folding angles are analyzed based on numerical simulation. The basic airframe parameters of SAW are listed as follows:\nBased on above airframe parameters, with the application of DATCOM, the 3D model of the SAW can be established and the flight status(\u03b4 = \u221230\u25e6 and \u03b4 = 60\u25e6) can be shown as follows:\n2020 Chinese Control And Decision Conference (CCDC 2020) 1817\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply.", + "With DATCOM symbols and the corresponding shape parameter given above, the related parameters in DATCOM such as SSPNE and SSPN corresponding to different folding angles can be obtained and listed below, based on which numeric simulation can be conducted to obtain aerodynamic performance of SAW under different folding angles:\nThus, substituting all the parameters and symbol values into DATCOM, the aerodynamic coefficients with folding angles ranging from \u221230\u25e6 to 60\u25e6 as well as control surface items can be obtained. Based on this, the aerodynamic performances under different folding angles with angle of attack ranging from \u22124\u25e6 to 10\u25e6 can be shown below:\nIt can be illustrated in Figure 5 that during morphing process, the lift coefficients appear linear correlation with angle of attack, meanwhile, it appears that with folding angle keeping at \u221260\u25e6, the lift coefficients have a greater slope, which shows folding upwards can make it easier to enhance altitude when the SAW takes off. Similarly, in Figure 6, with folding angle keeps at 30\u25e6, the drag coefficients are much smaller, which make it more suitable in cruise phase.\nTo clearly clarify the ability of SAW to conduct multimissions, an insightful description of the polar curves of the SAW under different folding angles are given:\nIn Figure 8, when \u03b4 = 0\u25e6, the lift-to-drag ratio reaches the maximum value, which means that keeping the wing-tip level is more suitable to conduct long-range cruise surveillance missions. Folding wing-tip leads to the decrease of lift-to-drag ratio, which makes the SAW have the high-speed airfoil with small-aspect-ratio and large-sweep-angle. With the wing-tip folding upwards, the SAW can dive fast and conduct high-speed sprint, which has a greater maneuverability.\n1818 2020 Chinese Control And Decision Conference (CCDC 2020)\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_80_0001471_j.engfailanal.2019.104223-Figure17-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001471_j.engfailanal.2019.104223-Figure17-1.png", + "caption": "Fig. 17. Gap between rear lower centre cowl and rear fender.", + "texts": [ + " In general, mounting holes of rear fender and rear cowl are oblong to accommodate frame deviations. Effect of rear fender in tolerance stack up analysis for minimum gap between rear lower centre cowl and rear fender is performed and it is found that 70% is contributed by assembly shift. The rear fender is tilted by 1.2 degree as shown in Fig. 16 due to assembly shift, thereby reducing the gap by 2mm. As per design, gap between rear cowl assembly and rear fender= h mm (without assembly shift) as shown in Fig. 17. The actual gap has been checked in tested vehicles and it is found that gap has reduced in all vehicles. With respect to actual gap measured in vehicle, the rear assembly along with frame is modelled as shell mesh with average mesh density as 2.5mm and is simulated for dynamic loading condition. Harmonic data used in electrodynamic shaker as shown in Fig. 3 is given as input in linear steady state dynamic simulation. Penetration of parts are noticed at a particular frequency. Fig. 18(a) shows displacement plot of assembly and Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002806_012079-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002806_012079-Figure1-1.png", + "caption": "Figure 1. mechanical drawing of roller brake tester [5]", + "texts": [ + " The roller brake tester measures the amount of axle braking force, the efficiency of the braking force and the percentage of imbalance braking force between the right brake and the left brake. The aim of this study is to analyse the vehicle braking force efficiency tested on the roller brake tester for three different modes of braking actions. BIS-ASE 2019 Journal of Physics: Conference Series 1517 (2020) 012079 IOP Publishing doi:10.1088/1742-6596/1517/1/012079 The mechanical drawing of roller brake tester is shown in Figure 1. It consists of main elements, rotating directions, forces and dimensions of the roller brake tester. The information presented in Figure 1 is very essential for understanding the working principle of brake tester system and for determination of the braking force. The appropriate position of the vehicle wheels placed on the brake roller tester is shown in Figure 2. When the brake roller tester runs, both rollers rotate at a constant speed (\u03c9W1 = \u03c9W2). These rollers are driven by a constant torque electromotor. When braking force is applied, the wheel decelerates and a torque MW is developed. The torque lever which is connected to the stator part of the electromotor deflects in the opposite direction from the rotations of rollers, electromotor rotor (\u03c9R), and vehicle wheel (\u03c9W)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure82.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure82.6-1.png", + "caption": "Fig. 82.6 SS GCI and AlSiC profile of von Mises stress (Model 3)", + "texts": [], + "surrounding_texts": [ + "In the coupled analysis, the thermal load was coupled with the structural load to find out the combined effect on brake disc models. The temperature induced at various time points was imported into static structural, and then structural loads were applied. The analysis was run for 36 s, i.e. the time taken by the vehicle to stop due to the application of the emergency brake. The output results of von Mises stress and total deformation developed in the model were recorded (Figs. 82.7, 82.8 and 82.9). Some important points that can be drawn from the analysis are: 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 979 \u2022 When compared to the Factor of Safety offered by the GCI models, the AlSiC models offer higher Factor of Safety. \u2022 For the same applied load, the AlSiC models have lower thermal stresses than the GCI models, as AlSiC material has greater thermal conductivity and heat dissipation capability. \u2022 The weight of AlSiC models is lesser when compared to the GCI models (GCI model having a weight of about 134 kg gets reduced to 54 kg in case of AlSiC material). 980 E. Madhusudhan Raju et al." + ] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure6.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure6.4-1.png", + "caption": "Fig. 6.4 Stage 3 of the proposed joining method", + "texts": [ + " Ganesh Narayanan Stage 2: In this stage, the external tube comes in contact with the punch and it also starts deforming above the internal tube. By the time, the internal tube keeps on deforming and acquires the shape of groove of the die. The process of stage 2 is shown in Fig. 6.3. Stage 3: As the upper portion of the die groove is heading inward, the leading edge of internal tube starts deforming in the inward direction. After a particular stage, the tubes get mechanically locked as shown in Fig. 6.4. Some experimental trials were conducted to demonstrate the joining method at lab scale. Figure 6.5 shows the experimental setup clamped in a universal testing machine. The tubes were displaced in the downward direction at a uniform cross-head speed of 1 mm/min. 6 Joining Concentric Tubes by End Forming: A Finite Element \u2026 69 The uniaxial tensile experimental tests were performed in order to study the mechanical properties of the tubes by means of universal testing machine. The specimen used as internal tube is mild steel tube of internal diameter 37" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002679_j.seppur.2020.117016-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002679_j.seppur.2020.117016-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of batch GAMS.", + "texts": [ + "30 g of FeCl2\u00b74H2O were dissolved in 200 mL deionized water under nitrogen gas with vigorous stirring at 85\u2103, and 25 mL of ammonium hydroxide (25%) was added rapidly into the solution. After 2 min, 50 mL of HA solution with the concentration of 0.02 g/mL was added. The reaction was permitted for 2 h under continuous nitrogen gas and stirring. The resultant black magnetic precipitates were washed several times with de-ionized water by magnetic decantation. The average size of these nanoparticles was about 10 nm by TEM characterization. GAMS experiments were done in a homemade batch gas-assisted magnetic separator shown in Fig. 1. It was made up of a nitrogen supply system, a glass flotation column with the inner diameter of 43 mm and length of 500 mm, and a hanging adjustable quadrate magnet with side lengths of 25 mm and field strength of 0.2 T. The typical experimental process was as following: Firstly, a certain volume of the feed nanoparticles solution in which magnetic nanoparticles had achieved the equilibrium adsorption for heavy metal was poured into the flotation column. Then, N2 with a given flow rate got through, and the hanging magnet down rapidly to the surface of feed solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003462_kem.861.182-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003462_kem.861.182-Figure2-1.png", + "caption": "Fig. 2 First-layer height [3]", + "texts": [], + "surrounding_texts": [ + "The first step in the study was to design the test specimens for determining tensile strength as per ISO 3167 A1 and to create a 3D model in SolidWorks as per the geometry and dimensions given in the standard." + ] + }, + { + "image_filename": "designv11_80_0002177_b978-0-12-803581-8.11760-6-Figure36-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002177_b978-0-12-803581-8.11760-6-Figure36-1.png", + "caption": "Fig. 36 Vertical target full-scale prototype manufactured by Plansee (high heat flux units) and Ansaldo Ricerche (support structure and integration).", + "texts": [ + " Active Metal Casting consists of casting a pure Cu layer onto a laser-textured and TiC-modified CFC surface138\u2013140 The laser-texturing enhances Cu infiltration into the CFC, and the TiC-modified CFC surface improves the wetting. The special laser treatment of the CFC surface produces a large number of closed conical holes (diameter B 50\u2013500 mm, depth 100\u2013750 mm) thus increasing the joined area and providing additional crack growth resistance. Due to the open porosity of the TiC modified CFC and laser machining, the cast Cu penetrates in the CFC up to 2 mm. An example of a full-scale component produced by Plansee, Austria, is in Fig. 36. AMC\u00ae was successfully applied both for flat tile and monoblock geometry. However, AMC\u00ae technology requires laser machining of CFC surfaces that might not be economically attractive for large-scale production. Laser induced stresses in the joined area and cracks induced during the joining process have been recently measured and modelled.122 The cross sections of the AMC from the CuCrZr to the Cu interlayer is shown in Fig. 37(a), with the laser structuring transition shown in Fig. 37(b) However, Plansee has recently improved AMC\u00ae by using silicon and titanium to modify the CFC surface (TiSi-AMC) (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003616_j.matpr.2020.08.536-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003616_j.matpr.2020.08.536-Figure10-1.png", + "caption": "Fig. 10. Directional Deformation.", + "texts": [ + "0 we have done the static structural analysis by fixing the spindle and given the dynamic load at track arm with the Triangular mesh. The results obtained are within the limits of factor of safety. Worm and Wheel Analysis: - The various tests done (Table 5) for the Worm and Wheel Gear: - Total Deformation Directional Deformation Shear Elastic Strain Von-mess Stress Fig. 9 explains the forces applied on the casing the amount of deformation occurred for a certain period of time causes total deformation. Fig. 10 explains the displacement occurs along the Fig. 9. Total Deformation. particular axis the directional deformation occurs. During damping there may be chances of multi axial stress acting on it with multiple stress components acting at the same time in the structure (Fig. 11). When the force applied on the casing it tends to cause deformation by slipping along planes parallel to the imposed stress. This causes deformation on the stress-imposed area (Fig. 12). The calculated data obtained with the help of the model simplifies the estimation of the payback period, life cycle, electrical parameters of the regenerative shock absorber control unit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure7-1.png", + "caption": "Figure 7. Stephenson1 mechanism.", + "texts": [], + "surrounding_texts": [ + "The primary objective of this study was to identify the joint that producesthe highest error in a mechanism due to clearance. Simulations were carried out for all mechanisms described in the previous section by introducing clearances in different joints, one at a time. The effect of different speeds (of crank) and clearance sizes was also studied. Some important results are discussed here. The results for the four bar mechanisms have been summarized in figures 13, 14 and 15. These show that Joint 3 is the most critical joint and the values vary significantly for 0.02 mm but they are very similar for 0.5 mm. So, sensitivity to speed is high for low clearance. Figures 16, 17 and 18 show mean deviation at different joints of different inversions of the Watt chain at 0.1 mm radial clearance. The trend of all the curves is the same. Joint 3 is the most critical joint in all the three mechanisms followed by joints 2, 4 and 5, i.e., the criticality order here is: Joint 3[ Joint 2[Joint 4[ Joint 5. Joint 6 in figures 5 and 6 is the grounded joint which is least critical of them all. Figures 19 and 20 show the criticality graphs for Stephenson1 mechanism and Stephenson1 mechanism with slider, respectively. Both the graphs show the same trend where the deviation is maximum at the output link and decreases as one proceeds to the other end of the mechanism. A common trend which seems to emerge is that the first joint next to the crank, which has a link connected to ground, is the most critical joint. After that the criticality ranking goes towards the crank and then goes further down to the joints in order of their distances from the crank. We would examine the validity of this hypothesis in the rest of the study. Figures 21 and 22 show the criticality graphs for Stephenson2 mechanism (1) and Stephenson2 mechanism (2). The order of criticality for both the mechanisms is: Joint4 [ Joint3 [ Joint2 [ Joint5. This is as per the hypothesis stated earlier. The most critical joint is the joint next to the crank, which has a link connected to the ground (i.e., Joint 4). The next most critical joints are the ones towards the crank, i.e., Joint 3 and Joint 2. Then comes the joints which are left, i.e., Joint 5. Joint 6 is the grounded joint. The two graphs look very similar as the structures of mechanisms are the same. The deviations increase significantly with increase in speed, if the clearance is at Joint 4. The values in first graph are slightly higher than the corresponding values in the second graph. The overall conclusion is that deviations are more, if the output link is closer to the crank. Similar results are seen in eight and ten-bar mechanisms as shown in figures 23 and 24. The joint next to the crank is the most critical joint. It should also be noted that the grounded joints also show a trend. The grounded joint in the first loop has a significant effect but rest of the grounded joints have very little effect on the output. Error in output generally increases with increase in the crank speed (expect for an aberration shown in figure 19). This happens for all clearance sizes. However, the change in mean deviation is quite small even with a 5 times increase in speed from 100 to 500 rpm. Figures 25 and 26 show the variation of mean deviation with increase in speed when a clearance of 0.1 mm is at Joint 2 and Joint 3, respectively. It can be noticed that the deviations usually increase with speed. This means that a machine working at higher speeds is expected to produce more error than a machine working at lower speeds. Increase in radial clearance size brings an almost linear increase in the deviation values as seen in figures 27 and 28. It is also worth noting from figure 27 here that the deviation of slider is nearly 50% for 0.1 mm clearance (0.0501 is 50% of 0.1) and it goes on increasing to 80% for 0.5 mm clearance (0.3977 is 80% of 0.5). This happens because at lower values of clearance, the number of collisions between journal and bearing is more. Due to more number of collisions, the journal is sent back to the center again and again, and thus it remains at the center for a larger amount of time as compared to the case of large clearances. Thus, in larger clearances, the journal appears to remain closer to the walls of the bearing for an extended period and spends less time going to the diametrically opposite wall." + ] + }, + { + "image_filename": "designv11_80_0001685_icems.2019.8921850-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001685_icems.2019.8921850-Figure6-1.png", + "caption": "Fig. 6. Distribution of magnetic flux density pre/post fault. (a) Normal. (b) 5- turn ITSC fault. (c) 15-turn ITSC fault. (d) 25-turn ITSC fault. (e) 35-turn ITSC fault. (f) 45-turn ITSC fault.", + "texts": [ + " Distribution of magnetic line of force pre/post fault is shown in Fig. 5. It can be found that there is an obvious difference in the distribution of the magnetic line of force before and after the fault, that is, the fault breaks the symmetry distribution of the magnetic field. Compared with the healthy motor, a higher saturation phenomenon occurs around the faulty slot which is caused by an additional magnetic field generated by the short circuit current. The magnetic flux density distribution of the motor pre/post fault under the rated load is shown in Fig. 6. The overall appearance of the magnetic field is symmetrically distributed in two poles in healthy condition. But in faulty conditions, due to the influence of the additional magnetic field which is generated by the short circuit, localized saturation occurs in the teeth around the faulty slot which is the first slot in Fig. 2(a) and the symmetrical distribution of the magnetic field are destroyed. In addition, this phenomenon deepens with the faulty severity increasing. IV. ELECTROMAGNETIC PERFORMANCES CALCULATION AND ANALYSIS The magnetic field is symmetrically distributed when the stator windings are symmetrical" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003694_978-3-030-58799-4_69-Figure17-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003694_978-3-030-58799-4_69-Figure17-1.png", + "caption": "Fig. 17. Courant number mean - arc-shaped case", + "texts": [ + " Apparently, the horizontal tailplane provides a negative lift when the elevator is not deflected, due to its initial negative incidence with respect to the aircraft centerline (ih < 0) and the downwash effect from the wing. Deflecting the elevator by \u221210\u25e6 increases the absolute value of the negative lift. On the other hand, deflecting the elevator by +10\u25e6 provides a positive lift, that is in absolute value lower than the previous case. The drag polar obtained for \u03b4e = \u221210\u25e6 is shown on the right side of Fig. 16. It is worth noting that the negative deflection of the elevator leads to a relevant increment of 78 drag counts for CD0. By looking at the drag breakdown, which is reported in Fig. 17, it is interesting to note CFD Prediction of Aircraft Control Surfaces Aerodynamics 85 Y cC l / c re f -1 -0.5 0 0.5 1 0 0.2 0.4 0.6 0.8 no deflection 10 deg deflection Fig. 11. Ailerons: spanwise wing loading at zero AoA with, and without ailerons deflection. CL C r 0.2 0.4 0.6 0.8 1 1.2 1.4 -0.045 -0.04 -0.035 -0.03 CFD VLM Fig. 12. Ailerons: rolling moment versus lift coefficient. that deflecting the elevator by \u221210\u25e6 (up) leads to higher drag and influences the vertical tailplane as well. This is due to the fact that the elevator leads to higher (negative) load in this case", + " NPB Job Requirements Job CPU Memory (MB) Runtime (sec) BT 2 1280 180 FT 2 5132 240 MG 4 26624 420 SP 6 5132 680 We run 110 jobs for the experiments: 34 BT, 34 FT, 25 MG and 17 SP jobs. The number of jobs with small requisitions was large, so the scheduling algorithms were able to perform the advance of a large number of jobs. Like in previous experiments, the cluster was used in a dedicated way and each algorithm was executed 30 times. For the BT, FT and MG jobs the standard input data was used. For the SP jobs, the number of iterations was changed to 160 and the problem size to 256 \u00d7 256 \u00d7 256. Figure 17 illustrates the makespan when using the four algorithms in this case. It is possible to observe that the EASY-backfilling algorithm presented improvement in execution time by 10% when compared to the SJF algorithm. It seems that EASY-backfilling algorithm respected the execution queue and executed the largest jobs in its intended time, unlike the SJF algorithm, which advanced smaller jobs and left larger jobs to the end. As a result, larger jobs end up delaying execution. Also, in this scenario, the FIFO achieved better performance when compared to the SJF, precisely by running the larger jobs in their time and not delaying them at the beginning" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000055_012044-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000055_012044-Figure1-1.png", + "caption": "Figure 1. Geometry and boundary conditions of numerical model without control.", + "texts": [ + " A frequency converter has been used to adjust the rotating speed. Windage power loss has been found to increase with increasing the rotating speed. In this study, numerical model has been applied to replicate the existing experiment. Geometry of the high speed gear as well as operating conditions of the numerical model has been made based on the existing experiment. The gear is characterized by a module of 5mm, a pressure angle of 20degree, a pitch diameter of 150mm and a face width of 24mm. A schematic of the numerical model without control is shown in figure 1. The air flow around high speed gear without control has been replicated by using the commercial software Altair HyperWorks AcuSolve. Pressure inlet and pressure outlet have been applied to the inlet and exit of the computational domains. The symmetry boundary condition has been applied to simulate the half computational domain. Unstructured mesh has been used in all regions. The grid number in computational domain is approximately 4 million. The air properties are characterized at the room temperature and atmosphere pressure" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure26.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure26.1-1.png", + "caption": "Fig. 26.1 Numerical modeling of orthogonal cutting", + "texts": [ + " Considering the effect of cutting edge on the cutting forces, temperature profile, the coefficient of friction and contact length, an attempt is made to understand the effect of cutting-edge radius on the cutting forces and temperature in the machining of difficult to machine alloy like Ti6Al4V. This is done by developing a numerical model for two cutting-edge radii of 20 and 40, and 60 lm for different cutting speeds. A finite element model using ABAQUS\u2122 is used for modeling and simulation of the tool and the workpiece. A 2D dynamic coupled temperature displacement model is used for simulating orthogonal cutting as shown in Fig. 26.1. 310 S. Mane et al. The element type used for both the tool and workpiece refers to CPE4RT elements. The tool is assumed to be a rigid body since there will not be any deformation in the tool. A reference point is located at the tip of the tool for providing velocity to the tool. 20, 40, and 60 lm are the three different cutting-edge radii of the tool used in the simulation. The workpiece used is a rectangular block of 3 5 15 mm. Finer mesh size is used near the tool\u2013workpiece contact region, as shown in Fig. 26.1. A total number of elements in the tool and the workpiece are 2392 and 39140, respectively. The cutting parameters for the modeling are listed in Table 26.1. The thermophysical properties of the tool and the workpiece are listed in Table 26.2. Since machining is a dynamic process, there will be high strains, strain rate, and temperature involved. To accommodate this, a Johnson\u2013Cook flow stress model was used for Ti alloy shown in Table 26.3. The Johnson\u2013Cook fracture model is given in Table 26.4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000029_s12239-019-0013-z-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000029_s12239-019-0013-z-Figure1-1.png", + "caption": "Figure 1. Wrenches associated with some links.", + "texts": [ + " The screw representation and the reciprocal screws of commonly used joints are listed in Table 1. 3.2. Screws and Reciprocal Screws of Links The wrenches associated with a link can be determined by finding the intersection of the systems of reciprocal screws associated with the joints of the link (Tsai, 1999), or by finding the common reciprocal screws to all the joint screws in the link (Zhao et al., 2009). Using the special cases of reciprocal screws discussed in Section 2.2, the wenches of some commonly used links are determined geometrically as shown in Figure 1, where a zero-pitch screw (pure rotation) or zero-pitch wrench (pure force) is represented by a line with a solid arrow and an infinitepitch screw (pure translation) or infinite-pitch wrench (pure couple) is depicted by a line with a hollow arrow. A general R-R link shown in Figure 1 (a) has two zeropitch joint screws $1 and $2 along the joint axes ua and ub. Hence, the reciprocal screws forms a fourth-order screw system. Let N be the common normal between the two joint screw axes $1 and $2. By case I, two zero-pitch force wrenches can be determined: W1 that intersects $1 and parallels to $2; W2 that intersects $2 and parallels to $1. By case I or II, a zero-pitch force wrench W3 that intersects both $1 and $2 at right angles can be found. By case III, an infinite-pitch couple wrench W4 which is perpendicular to both $1 and $2 can be determined. An R-P link shown in Figure 1 (b) has one infinite-pitch joint screw $1 along the joint axis ua and one zero-pitch joint screw $2 along the joint axis ub. By cases I and III, two zero-pitch force wrenches W1 and W2, both intersecting $2 and perpendicular to $1, can be determined. By case III, two infinite-pitch couple wrenches W3 and W4, both perpendicular to $2 and reciprocal to $1 by case IV can be determined. In Figures 1 (c) to (f), the joint screws are not drawn for clarity. An R-S link shown in Figure 1 (c) has one zeropitch joint screw along the axis of the R joint ub and three zero-pitch joint screws along the principal axes whose origin is at the center of the S joint. Let N be the line normal to ub that passes through the center of the S joint. By case I, two zero-pitch force wrenches W1 and W2 which are intersecting or parallel to the four joint screws can be determined. An S-C link, whose C joint axis passes through the center of the S joint, shown in Figure 1 (d) has five joint screws, however, they are not all independent: the zeropitch joint screw along the C joint axis and the zero-pitch joint screw along the same axis of the S joint are coaxial, hence they are dependent and introduce a passive degree of freedom that allows the intermediate link to rotate freely about the axis of the C joint. Hence, one of the two zeropitch wrenches can be eliminated. Then, by cases I and III, two orthogonal zero-pitch force wrenches W1 and W2 which are intersecting or parallel to the three zero-pitch joint screws and perpendicular to the infinite-pitch joint screw along the C joint axis can be determined", + " In this section, using the method introduced in Section 3 to find the reciprocal screws of links, five independent wrenches acting on a rigid body of a planar mechanism are identified, and then the twist of the body is found graphically using the reciprocity of screws. A planar one degree of freedom four-bar mechanism is shown in Figure 2 whose rotation axes of R joints are parallel to each other and pointing along the z-axis. The coupler link, bce, is connected by two R-R links, ab and dc. Since each R-R link exerts four wrenches as shown in Figure 1 (a), the coupler seems to be acted upon by eight wrenches, but they are not all independent. Using the list of the dependency of screws in various cases classified by Huang et al. (2013), the dependent wrenches can be eliminated as follows. Four wrenches associated with the R-R link, ab, of the planar four bar in Figure 2 are two parallel force wrenches W1 and W2 acting along the z-axis at the R joints labeled a and b, one couple wrench (not shown in Figure 2), and one force wrench W3 along the line joining the two R joints at a and b", + " For this, five independent wrenches acting on the wheel hub by the connected links for the bump-rebound motion and steering motion are determined, then the instant screw axis and pitch of the wheel for each motion can be obtained by the twist reciprocal to the five wrenches using Equation (9). 5.1. Instant Screw Axis of Bump-rebound Motion Figure 4 shows an RSSR-SS spatial mechanism, which is the kinematic representation of the double wishbone type suspension. The wheel hub is connected to the vehicle body by two R-S links, each of which exerts two zero-pitch force wrenches on the wheel hub acting at the center of the S joint as shown in Figure 1 (c), and by an S-S link which has a zero-pitch force wrench acts through the centers of two S joints as shown in Figure 1 (e). These five wrenches are independent, and the twist reciprocal to the five wrenches determines the instant screw axis of the wheel with respect to the vehicle body for the bump-rebound motion and its pitch. Figure 5 shows an RSSS-SC spatial mechanism which is equivalent to the McPherson strut type suspension where the wheel hub is connected by an S-C link which exerts two force wrenches acting perpendicular to the S-C link at the S joint, by an S-S link with a zero-pitch force wrench acting along the S-S link, and by an R-S link with two zero-pitch force wrenches" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000728_j.oceaneng.2019.106199-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000728_j.oceaneng.2019.106199-Figure1-1.png", + "caption": "Fig. 1. Body-fixed coordinate frame A and global coordinate frame U.", + "texts": [ + " In Section 3 the design methodology and the convergence to zero of the tracking errors are presented. In Section 4, a method to controller tuning is shown and two tests are implemented with the purpose of checking the effectiveness of the new control law. Main conclusions of the research are described in Section 5. To determine the complete configuration of the marine vessel (position and orientation) six independent coordinates (surge, sway, heave, roll, pitch and yaw) are required, as can be seen in Fig. 1. However, these six degrees of freedom are commonly reduced to three. This is done because the remaining degrees are open loop stable for most marine vessels. Consider the ship model described in Serrano et al. (2014); B\u00f8rhaug et al. (2011), _x \u00bc ucos\u00f0\u03c8\u00de vsin\u00f0\u03c8\u00de _y \u00bc usin\u00f0\u03c8\u00de \u00fe vcos\u00f0\u03c8\u00de _\u03c8 \u00bc r (1) In (1) n \u00bc \u00bdx; y;\u03c8 \ufffdT denotes the position vector in the earth-fixed reference frame, (2) \u03c5 \u00bc \u00bdu; v; r\ufffdT 2 R3 represents the velocity vector in the body-fixed reference frame, M, C\u00f0\u03c5\u00de and D vessel inertia matrix, centrifugal and coriolis matrix and the hydrodynamic damping matrix, respectively; \u03c4 \u00bc \u00bdTu;Tr\ufffd T is the control input vector, where Tu and Tr are the surge and yaw control, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000789_isie.2019.8781273-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000789_isie.2019.8781273-Figure1-1.png", + "caption": "Fig. 1. The structural of NDUV", + "texts": [ + " The proposed controller has a faster response, a smaller steady error and overshoot. The interference in the water is considered. The simulation and experiment results demonstrated that the controller runs more stable than the conventional PID control law. (3)The dynamic model of NDUV is explored. Six-freedom dynamic model is established elementarily. Depth and attitude control system model are obtained. The simulation and experiment results prove that the explored model is valid. The structure of NDUV is shown in Fig. 1. It is mainly composed of a sphere hull and four propellers. The two hemispherical shells comprising the sphere hull are connected by a flange. The four low-cost motors with rotor-like propellers are respectively mounted on the arms of the flange plate. The control system, including a cheap inertial measurement 978-1-7281-3666-0/19/$31.00 \u00a92019 IEEE 1121 unit (IMU), a low-cost depth sensor, wireless communication module and a civil lithium battery, is placed inside the hull. When NDUV dives into the water, motor 1 and motor 3 rotating in opposite directions to motor 2 and motor 4 generate the upward thrust and push NDUV into the water" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001335_chicc.2019.8865877-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001335_chicc.2019.8865877-Figure2-1.png", + "caption": "Fig. 2: Models of directed sensor", + "texts": [ + " Remark 1: Avoiding obstacle based on artificial potential function method is widely used. However, it is easy to cause local minima that makes the mobile sensor unable to reach the intended position. At the same time, the trajectories in the process of obstacle avoidance are not smooth enough, which causes a lot of energy waste. In order to solve the problem, we design a switching strategy with the potential function to avoid local minima in this paper, based on stream function. Remark 2: It is believed that there are two orientation sensor models, as shown in Fig.2, where type (a) is called model A; type (b) is called model B. In model A, the perception area is fan with fixed direction, it can be expressed with Si1 (qi, R, \u03b5, \u03b2). In model B, there is a sectorial perception within its sensing radius. it can work in several directions around it and can be expressed with Si2 (qi, R, \u03b5, \u03b2, \u03c9). Compared with model A, model B needs less sensors, therefore we choose the model B in this paper. Remark 3: It is generally believed that 1-coverage topology is the most energy-efficient structure, it is proven in literature [9], as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002950_042044-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002950_042044-Figure5-1.png", + "caption": "Fig 5. Finite element mesh generation of temperature field", + "texts": [ + "2) Where, h0 is the natural heat dissipation coefficient of the object in the case of still air, v is the relative speed between the air and the surface of the object, unit is m/s, k is the influence coefficient of the heat dissipation efficiency in consideration of the air flow blowing. From the material of the permanent magnet coupler, h0=14W/(m2\u00b7K), k is 0.5, and the heat dissipation coefficient is calculated according to the linear speed of each part, unit is W/(m2\u00b7K). Simplifying the three-dimensional model of the permanent magnet coupling to established the finite element analysis mode, as shown in the Fig. 5. The conductor cylinder and the outer steel plate on the same side are set as the thermal load, assuming that the temperature distribution of the conductor cylinder and the outer steel plate is uniform, add internal h load to each generation and the surface heat transfer coefficient are meshed and analysed. Fig. 6 is the simulation analysis result of temperature field of cylinder permanent magnet prototype, and under 20rpm normal temperature. It can be seen from Fig.6 that the temperature of conductor cylinder, inner aluminium yoke and permanent magnet is higher, and the temperature of outer aluminium yoke, permanent magnet and steel frame is lower" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002288_s13369-020-04439-0-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002288_s13369-020-04439-0-Figure15-1.png", + "caption": "Fig. 15 Modeling of assembly defect", + "texts": [ + " This theoretical gap shape E can be defined by: E r12 + r21. (20) where r12 and r21 represent the pitch radius of the gear 12 and 21. Practically, it is very difficult to verify this assembly condition. In this section, a defect of distance between the wheel axles noted by the algebraic value (a) is proposed. Thereafter, it can be supposed that that the second transmission shaft (gear 21) is shifted from its theoretical position inducing an assembly defect. The modeling of this defect is shown in Fig. 15. Teeth gap provokes the average stiffness value kmoy and therefore the evolution of the gear mesh stiffness since the pressure angle will be changed. The new pressure angle can be expressed as: \u03b1\u2032 cos\u22121 (rb12 + rb21 E + a ) . (21) The assembly defect can change also the primitive radius and the tooth step. The new primitive radii are written by: r \u2032 12 rb12 cos(\u03b1\u2032) and r \u2032 21 E + a \u2212 r \u2032 12. (22) A change in this parameters leads to a modification in the contact ratio that will be affecting the dynamic behavior of the system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003110_s40194-020-00955-7-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003110_s40194-020-00955-7-Figure5-1.png", + "caption": "Fig. 5 Particle concentration distribution of powder flux. (a) Initial state. (b) Adjustment of defocusing distance. (c) Adjustment of powder feeding angle. (d) Relationship of nozzle position", + "texts": [ + " 7 of [8], powder transport ratio had been measured only after adjusting the defocusing distance and powder feeding angle respectively, which ranged from \u2212 2 to 10 mm by 2 mm increments for \u03c7, and from \u2212 3\u00b0 to 3\u00b0 by 0.5\u00b0increments for \u03b1. By weighing the powder quantity stored between above two steel sheets always set at 3 mm distance, the experimental data could be estimated and distributed in Fig. 6. Similar to the experiment in Section 3.1 before, adjustment of defocusing distance \u03c7 (+ 4 mm) and powder feeding angle \u03b1 (+ 3\u00b0) was substituted into the theoretical model successively. In Fig. 5(a\u2013c), the concentration distribution of powder particles calculated by Eq. 29 is projected into a two-dimensional Table 3 Coordinate information of key points in powder flux Initial state \u2192Defocusing distance +\u03c7.(\u03c7 = 4 mm) \u2192Powder feeding angle +\u03b1 (\u03b1 = 3\u00b0) A P2 P1 O A1 P2 P1 M A2 P2 P1 N Experimental data (\u2212 5.2, 8.9) \u2212 2.11 2.0 \u2212 0.1 (\u2212 5.2, 12.9) \u2212 1.5 4.0 1.3 (\u2212 9.9, 10.3) \u2212 7.1 \u2212 2.3 \u2212 4.7 Theoretical data (\u2212 5.2, 8.9) \u2212 2.0 2.0 0.0 (\u2212 5.2, 12.9) \u2212 1.0 3.9 1.5 (\u2212 9.9, 10.3) \u2212 6.7 \u2212 2.4 \u2212 4.6 Figure 5(a) Figure 5(b) Figure 5(c) A, A1, A2\u2014outflow position, P2\u2014left boundary of powder flow, P1\u2014right boundary of powder flow, O, M. N\u2014injection position space of 13.8 \u00d7 12.9 mm2 (\u2212 9.9 mm \u2264X \u2264 3.9 mm, 0 \u2264 Z \u2264 12.9 mm), which can be illustrated in Fig. 5(d). According to the cloud scale, the powder flux conforms to a Gaussian distribution and its key point coordinates are extracted in Table 3 for comparison with the experimental measurements. As the model is solved analytically, the calculation of the outflow position (A, A1, A2) is accurate, while small deviations for the injection position (O, M, N) and boundary of powder flow (P1, P2) exist probably due to inaccurate substitution of initial conditions, such as \u03b3 and \u03c6. Despite some errors, the results show that this model can truly reflect the distribution and spatial position of the powder flux after adjusting the powder injection parameters", + " The change of the powder flux position relative to the fixed molten pool causes a different \u03bePT, which is discussed further. Moreover, in the theoretical calculation process, the powder feeding rate Vm is set as 2.322 g/min, which was measured from the actual lateral powder feeding nozzle in laboratory. In a three-dimensional space, the real powder stream can be intercepted by multiple planes parallel to the XOZ direction (Fig. 1), and then, slice photos containing particle distribution information are stacked, which can be regarded as the particle concentration distribution in the two-dimensional space shown in Fig. 5. The shape of the molten pool can also be projected on the XOZ plane to obtain the molten pool width. Finally, the powder efficiency can be estimated by mass ratio of powder particles falling within this width range to all powders injected from the nozzle. In essence, it is a dimension reduction operation from 3D to 2D. For the experiment as described in Section 3.2, the onedimensional variation of \u03bePT with \u03c7 or \u03b1 is calculated by Eq. 29 and plotted in Fig. 6, in which the experimental data under 3 mm molten pool width condition was also distributed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002189_acit47987.2019.8991028-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002189_acit47987.2019.8991028-Figure3-1.png", + "caption": "FIGURE 3. Work principle of plate-inclined plunger pump/motor.", + "texts": [ + " Results showed that the simulation model was corrected and the control strategy can improve the EHHT\u2019s control performance. II. PRINCIPLE ANALYSIS OF EHHT Electro-hydraulic servo plate-inclined plunger hydraulic transformer is a swash plate plunger-type hydraulic component, with several advantages such as compact structure, small radial size, low mass, low moment of inertia and so on. It achieves the function of oil suction and discharge through the volume change generated by the reciprocating movement of the piston against the cylinder bore. The schematic diagram is shown in Fig.3. As there is an angle \u03b9 in the swash plate, when the cylinder is rotated, the plunger will reciprocate in the cylinder cavity. The displacement equation is s = R tan \u03b9(1\u2212 cos\u03d5) (3) Where s is the displacement of plunger at the top dead center (m), R is the radius of plunger distribution circle (m), \u03b9 is the tilt angle of swash plate (\u25e6), \u03d5 is the angle of plunger versus the top dead center (\u25e6). Swash plate piston pump / motor is featured by compact structure, small radial size, low mass and low moment of inertia" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001455_012165-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001455_012165-Figure3-1.png", + "caption": "Figure 3. 3st Principal stress ( Max: -29.73 MPa)", + "texts": [], + "surrounding_texts": [ + "IOP Conf. Series: Earth and Environmental Science 343 (2019) 012165 IOP Publishing doi:10.1088/1755-1315/343/1/012165" + ] + }, + { + "image_filename": "designv11_80_0001303_chicc.2019.8865927-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001303_chicc.2019.8865927-Figure5-1.png", + "caption": "Fig. 5 A potential function analysis between two UAV s", + "texts": [], + "surrounding_texts": [ + "Set the translation vector describing the position of UAV i as: = i i i x y (15) 1) Potential energy of interaction and acting force As shown in the figure below, each UAV in the formation aims to reach its target position ( ) ig . Therefore, for the UAV i, the distance between its actual position and the target position should be considered: iig g iL (16) For the sake of convenience, the target error is obtained by ig igL L . The distance between the UAV i and the UAV j is defined as ij j iL , and the distance error is as ij ijd ijL L L , where ijdL is the expected distance between the two UAVs. The artificial potential energy functions between the UAV i and the UAV j and relative to the target position are: 2 2 1 2 1 2 ij f ij ig f ig U k L U k L (17) Where fk is the gain constant of the formation. The total structural potential energy from all UAV s at the UAV i is: n i ij ig j i U U U (18) The corresponding structural force can be obtained from the negative gradient of the structural potential energy, so the overall structural gravity ,att iF acting on the UAV i can be acquired from the resultant force of the target point and other UAVs on the UAV i: , n att i i ij ig j i n j i g i f ij f ig j i ij ij F U f f k L k L L L ! (19) Where, n represents the number of UAVs in formation, and the time derivative of ,att iF can be expressed as: , ( ) n n att i f j i f j f j f j j i j i F k k k n k (20) 2) Air collision prevention The attraction function between two UAVs stipulates that, in order to maintain the desired distance between UAVs, in the case that ij ijdL L\" , the repulsion force should be exerted on the UAVs. However, under the action of gravity near the target and other UAV attraction functions, it is possible that the repulsion function between two UAVs fails to avoid collisions. So the strategy is to add a counteracting force of the same magnitude. In order to avoid inter-aircraft collision, the repulsive force between UAVs is defined as follows: , 0 , 0 , 0, i j att i ij ijr ij ij F L L Lf L L # (21) The repulsive force is the same as the gravitational force and in the opposite direction. 0L represents the safe distance where the collision will not occur, therefore, the distance between the two UAVs should be greater than 0L . 3) Obstacle avoidance Assume that the environment is static and the obstacle position is fixed. In order to realize obstacle avoidance, repulsive potential energy must be defined, and which will only be valid in a region near the obstacle. For UAV i near the obstacle k , the repulsive force can be expressed as: , , 1 0 , 0 , 0, n i ok att i r ij iok j ir iok iok iok F f b L L f L L L $ % # & '& ' ( ) (22) Where 1b is the minimum boundary in the nested saturation control, iok i okL corresponds to the distance between the center of mass of the UAV and the nearest boundary of the obstacle k , ok is the position of the obstacle, and 0L is the distance at which the repulsion function starts to operate." + ] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-Figure1-1.png", + "caption": "Fig. 1. Concept of tendon-driven elastic telescopic arm", + "texts": [ + " In addition, Spiral Zipper [9] that forms a cylindrical shape by shaping a long, thin plastic band upward into a spiral, or a carbon fiber reinforced plastic (CFRP) convex tape [10] have been proposed. Both methods achieve a high extension-to-contraction ratio by winding the band, and the Spiral Zipper achieves the ratio of over 14: 1 and control the tip position. However, in both methods, Young\u2019s modulus of the arm structural material tends to be small and buckling, so there is a problem that the increase of the weight due to the elongation can not be tolerated. From such a background, we have proposed the Tendondriven Elastic Telescopic Arm [11] as shown in Fig. 1 as a small diameter arm that is capable of linear motion and bending motion in order to achieve a lightweight, compact, 978-1-7281-6667-4/20/$31.00 \u00a92020 IEEE 1328 actuator saving, and a high extension-to-contraction ratio. In our previous work, we separately confirmed and demonstrated the mechanical feasibility of both motions. In this paper, we integrate the linear motion and bending motion. We propose the linear mechanism that can extend and contract one node at a time by a linear movement mechanism using a slide screw, and clarify its effectiveness by experiments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002071_iccas47443.2019.8971752-FigureI-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002071_iccas47443.2019.8971752-FigureI-1.png", + "caption": "Fig. I (a) For a stand-riding type WIP, a rider moves its weight by rotating a whole-bod y around an ankle, (b) for a sit-riding type WIP wheelchair, a rider rotates an upper-body around a waist.", + "texts": [ + "eywords: Wheeled Inverted Pendulum, Sit-ridin g type, Active slider, Wheelchair 1. INTRODUCTION control of two-whe eled wheelchair system. For a stand-riding type WIP personal mobility vehi cle[13][14][15] as the SEGWAY, a rider moves its body weight back and forth by rotating a whole-body around an ankle as shown in Fig. I(a). This make it possible to easily accelerate and decelerate the vehicle and quickly react against its driving behavior. However, for WIP wheelchairs, a rider operates it in a sitting posture and a weight-shifting motion is achieved by only rotating an upper-body around a waist as shown in Fig. I(b). This causes a drop of acceleration and deceleration ability of the wheelchairs because the movement of rider 's CoG is restricted than that in a standing posture. To realize large acceleration and deceleration, a rider should largely bend a waist back and forth , which makes a rider feel uncom fortabl e. As more, a rider in a sitting posture feels anx ious when the wheelchair incline s back and forth while accelerating and deceleratin g[ 16]. To cope with the problem, a slider mechanism has been introdu ced to impro ve driving performance and to 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002873_icpes47639.2019.9105570-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002873_icpes47639.2019.9105570-Figure12-1.png", + "caption": "Fig. 12: Flux- and loss distribution of the fully-pitched integer-slot winding with q = 1. Results are listed in Table II.", + "texts": [ + " The considered winding layouts are listed in Table A-I. 2) Simulation and Results: The influences of the harmonics on the losses are shown in Fig. 11. Comparison of the fully-pitched windings (W/\u03c4p = 1) shows that reducing the amplitude of the harmonics yields a significant reduction of the rotor and PM losses. Only the stator losses are nearly constant for all considered SPP, which is in good agreement with the results of the harmonic analysis in Fig. 7. The loss distribution and the flux distribution as well as the flux lines are presented in Fig. 12. It can be seen, that when using integer-slot windings the stator is the main source of the machine losses. Additionally, the effects of chording on the losses can be seen in Fig. 11. Compared to the analysis before, shortening the coil-span has directly an effect on the stator losses due to the reduction of the fundamental component of the MMF (compare Fig. 10). The rotor and PM losses are nearly on the same level, which can also be seen in Table II, where the values of the loss calulation of the fully- and fractionalpitched windings are listed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001227_rusautocon.2019.8867687-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001227_rusautocon.2019.8867687-Figure4-1.png", + "caption": "Fig. 4. Calculation of a crane trolley coordinate on the x-axis", + "texts": [ + " 3 shows a diagram consisting of a camera matrix and a lens aperture. Knowing the physical dimension Z of an object and the Z' dimensions on a digital camera matrix it is possible to calculate the distance to object y by expression y = Z\u00b7f/ Z' (1) where f indicates a camera focal length. The y coordinate is determined by the similarity of triangles. The area of the label rectangle is used as the measured parameter Z. Using the above feature of similarity of triangles and the obtained value y we can calculate the position of the crane on the x-axis (Fig. 4). The position of the trolley must be calculated from the center of the coordinates since the distance from the camera increases the viewing area. The formula to calculate the x coordinate is following X = Zx \u2018\u00b7y/f (2) where x \u2013 distance from the camera axis to label center; Zx ' - distance from the camera axis to label center on matrix; y \u2013 trolley coordinate on y-axis; f - focal length. As a result, the movement transients of the bridge crane trolley position in two axes are obtained. To determine the speed, it is necessary to differentiate the resulting array of coordinates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure5-1.png", + "caption": "Figure 5.Results of top gating system. Filling time(a) illustrate the uneven of liquid. Combined defect parameter(b) and Probabilities defect parameter(c) show the location of shrinkage defects or gas porosity.", + "texts": [ + " Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 After the optimization of gating system, the selected solution will be printed for application in reality experiment. Material for 3D printing is PLA and the final forming metal is ZL104 which the temperature for casting is 750\u2103. Thermo stability plaster is employed as the material of casting shell due to its advanced liquidity, low thermo conductivity and high accuracy in re-model. The results of top gating system are shown in Figure 5. The Fig explains the state of metal fluid that arrive the vault of mode cavity in 1.6sec. On account of the fluid dropping into the cavity directly, speed of fluid is fast as purple area. Thus the surface of fluid is unsmooth which means the appearance of turbulence. The final results of filling can be explained by Figure 5(b) and Figure 5(c). The combined defect parameter analyses the possibility of defect while the probabilistic defect parameter proof that the shrinkage defects and gas porosity will occur between the gating system and the impeller like the zone printed in yellow. In bottom gating system, the metal fluid is poured through the bottom of cavity. The solidification time is shown as Figure 6, which means that the solidification of impeller is in an orderless situation. Many defects located in the blades according to the combined defect parameter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure24.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure24.4-1.png", + "caption": "Fig. 24.4 Definition of SPR parameters", + "texts": [ + " After the completion of each simulation, temperature at an element that is in contact with the rivet tip was evaluated with respect to punch displacement. 24 Temperature Prediction During Self-pierce Riveting \u2026 287 The SPR parameters considered for the ANN modelling are Coulomb\u2019s coefficience of friction (\u00b5), die tip height (Dh), die depth (Dd), rivet length (Rl), upper sheet thickness (Ut), lower sheet thickness (Lt), upper sheet K (UK), upper sheet n (Un), lower sheet K (LK), lower sheet n (Ln). Here, K and n are the strength coefficient and strain hardening exponent of the sheet materials. The parameters are defined in Fig. 24.4. The levels for these parameters (shown in Table 24.2) were selected based on available SPR data and material properties. The range was chosen to satisfy the practical range available for these parameters. In the present work, ANN model was developed to predict the temperature evolution during SPR. A total of 440 FE simulations were carried out for ten parameters (Table 24.2) with several levels for each parameter. To predict the temperature evolution during SPR, the temperature-displacement curves from FE simulations were divided into segments at equal intervals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001707_icems.2019.8921756-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001707_icems.2019.8921756-Figure7-1.png", + "caption": "Fig. 7. Built-in damped sandwich structure", + "texts": [ + " The damped sandwich structure is generally divided into two types, a built-in damped sandwich structure and an external damped sandwich structure. Embedding the damped sandwich into the stator core to reduce the natural frequency for the purpose of weakening the electromagnetic vibration The mass of the damped sandwich is 1300 kg/m3, the elastic modulus is 7.86MPA, the poisson's ratio is 0.47. The damped sandwich has an inner diameter of 180 mm and an outer diameter of 230 mm. The built-in damped sandwich structure is shown in Fig. 7. The external damped sandwich stator core is embedded in t the core, and the sandwich is embedded on both sides of the stator core. The thickness of the core is still 117mm. The parameters of the damped sandwich are: outer diameter is 230 mm, inner diameter d is 180 mm, thickness is 15 mm, and total thickness is 30 mm. The external damped sandwich structure is shown in Fig. 8 The built-in damped sandwiches are embedded in the middle of the stator core. The thickness of the damped sandwich is 10mm, 15mm, 20mm, 25mm, 30mm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003106_mrc.2019.153-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003106_mrc.2019.153-Figure9-1.png", + "caption": "Figure 9. (a) Deformation and failure of the FBCCY hybrid structure under uniaxial tensile loading and (b) deformation and failure of the FBCCY hybrid structure under pure bending loading. The zoom-up views showing the failure at the joints in the tubes (marked by frames).", + "texts": [ + " East Carolina University, on 09 Dec 2019 at 03:52:12, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. deformation and failure behaviors of the FBCCY hybrid structure under pure bending loading. In the model, the plate thickness h1 is 1000 nm, the thickness of nanolattice tube t is 50 nm, and the number of unit cells n along the thickness of the structure is 4. For comparison, we have also analyzed the stress distribution for the FBCCY hybrid structure under uniaxial tensile loading. Figure 9(a) shows the stress distribution for the FBCCY hybrid structure under uniaxial tensile loading. We find that the maximal von Mises stress occurs at the joints (at the tube side) between the plates and the lattice, which are marked by frames in the zoom-up view in Fig. 9(a). Figure 9(b) shows the stress distribution for the FBCCY structure under pure bending loading. We find that the maximal von Mises stress also occurs at the joints (at the tube side) between the plate (under tension) and the lattice, which are also marked by frames in the zoom-up view in Fig. 9(b). It is seen interestingly that the stress distribution at the joints under uniaxial tensile loading is very similar to the stress distribution at the joints (with the plate under tension) under pure bending loading. Hence, the failure behavior under pure bending loading shares similarity with that under the uniaxial tensile loading. Hence, the conclusions obtained under uniaxial tensile loading are likely to be also applicable to those under pure bending. Clearly, more work is needed to study the failure behavior of the hybrid structures under complex loading conditions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001105_j.jnnfm.2019.104165-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001105_j.jnnfm.2019.104165-Figure3-1.png", + "caption": "Fig. 3. Schematic diagram of idealised steady state Kaye effect, and the coordinates used in our control volume analysis.", + "texts": [ + " A viscoelastic fluid subjected to a non-zero train rate for a finite period of time will continue to carry a stress for ome time (characterised by the relaxation time) after the strain rate eturns to zero. In the following we consider a jet of incompressible iscoelastic liquid with a relaxation time \ud835\udf06. A Newtonian fluid is repreented by \ud835\udf06 = 0 . A f h i \ud835\udf19 f t c T a F ( W l t t v v ( t i t T \ud835\udf03 t d t c t a s s \u0394 i s r \u0394 a \ud835\udc6d W t \u0394 C \ud835\udc6d 3 r d U \ud835\udf49 w \u2207 \ud835\udf49 a o a t t u e i \ud835\udc37 Consider a jet with constant flow rate undergoing the Kaye effect. schematic diagram of this is given in Fig. 3 . The primary jet emerges rom a reservoir with orifice of radius R 0 , and falls a height H before itting the heap and bending. The centreline of the jet round the bend s assumed to have a constant radius of curvature R , through an angle . The system is assumed to be in a quasi-steady state: the timescale or changes in the geometry of the flow is significantly greater than the imescale for fluid to pass from the reservoir to the secondary jet. The oordinate system x, y has origin at the centre of the reservoir orifice" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001815_s12206-019-1106-3-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001815_s12206-019-1106-3-Figure5-1.png", + "caption": "Fig. 5. A schematic representation of three different slicing methods for part build: (a) Uni-directional slicing; (b) multi-axis slicing; (c) curved layer slicing [85].", + "texts": [ + " In fact, surface inaccuracy of AM part, the notorious and inevitable disadvantage of all AM parts for a long time, is occurred mainly due to the planar layer itself [81]; It does not matter whether those are made from either STL models [82] or free form solid models [83]. It actually even does not matter whether the layer is generated either by using adaptive slicing or by using uniform slicing as well [84]. Selection of the type of a layer geometry, whether it is planar or curved while slicing the designed part, can be thereby an additional way of optimizing an AMM process according to a current research work as shown in Fig. 5 [85]. Kalmanovich (1996) has first introduced non-planar or curved layer models for laminated object modeling (LOM) process [86] in order to acquire additional strength and reduce build time for the part build. Using \u2018height grid\u2019 representation those layers are bonded together as multiple non-planar surfaces originally on non-planar base geometry. Klosterman et al. (1999) [87] also developed a curved layer LOM machine using ceramics and composites with increased build speed and less stair step effect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003718_icra40945.2020.9197100-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003718_icra40945.2020.9197100-Figure7-1.png", + "caption": "Fig. 7: Maximum grasp strength on different surfaces. Different materials compare the effects of friction, compliance, and adhesion. *These trials were stopped at 30 N to protect the PLA links.", + "texts": [ + " Furthermore, the gripper typically did not exert large localized forces onto the object even at high cable tensions; during the pull test the force in the grip cables peaked at nearly 110 N, but the maximum inward grip force did not exceed 2 N. Grasp success depended on surface properties such as friction, system compliance, and use of adhesives such as magnets and gecko-inspired films. We tested four different interfaces: a gecko-inspired adhesive [25], soft silicone foam, the PLA link without an interface, and magnetic tiles (Fig. 7 top, from left to right). To further study these effects, we tested these interfaces on four distinct geometries: a cylinder, dish, rectangle, and an irregularly shaped mirror mount (Fig. 7 bottom, from left to right). A force gauge (Mark 10) measured the maximum pull force the gripper held on each object. The gripper was tested to a maximum of 30 N to avoid damaging the 3D printed linkages. The best performance resulted from the gecko-inspired adhesive interfaces, which reached the 30 N limit on all surfaces. Our gripper performed similarly on curved surfaces to prior gecko-adhesive grippers [8], [10], [22], [23] with two important differences. 1) The joints in this gripper rotated independently and wrapped more closely around objects with protrusions, improving the adhesive force in arbitrary configurations", + " Highly under-actuated grippers offer a unique set of tradeoffs. These grippers can deploy to long lengths for grasping large objects while still maintaining conformation to arbitrary shapes. These grippers perform best when wrapped around an object, and struggle with pinch grasps and inhand manipulation. Mechanical complexity of the system is high relative to parallel jaw grippers or grippers made from soft elastomers. Highly under-actuated grippers can also conform to concave surfaces through the process of sequential deployment (Fig. 7 - Dish). These grippers can be used on or near humans with relative safety because they grasp with low, distributed, pressure (Fig. 9-A) and can even follow along a non-continuous path (Fig. 9-B). Finally, using dynamic actuation or a pair of linkages, these grippers can capture non-adjacent objects (Fig. 9-C). Highly under-actuated linkages blur distinctions between grippers comprised of low-DOF linkages and soft continuums. There are lessons and designs from these technologies that can flow in both directions" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003739_s12239-020-0106-8-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003739_s12239-020-0106-8-Figure4-1.png", + "caption": "Figure 4. Schematic diagram about force analysis of one petal-shaped connecting block in contact with the ground.", + "texts": [ + " Assuming that the acceleration and angular acceleration of the tire under driving state are a and \u03b1 respectively, then the mechanical equation of the wheel can be constructed: (1) where T is the supporting force of the ground to the tire; m is the quality of the tire; m is the mass of the tire; Fx is the horizontal force acting on the tire by the axle under the driving condition; J is the inertia moment of the tire; R is the moment arm of driving force; Mf is the rolling resistance moment. As shown in Figure 4, when there is only one petalshaped connecting block in contact with the ground, the ground will give the petal-shaped connecting block at the interface a vertical upward supporting force T due to the impact of the car body load. At this time, the two springs hinged on the other side of the petal-type connecting block and the piston fixedly connected with the petal-type connecting block through bolts that is subjected to the gas in the cylindrical cavity and annular cavity respectively m J R d x d d f T = H +G a = f - F \u03b1 = M - f -M i i i provide the opposite forces of , and opposite pressure to the petal-type connecting block" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure82.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure82.1-1.png", + "caption": "Fig. 82.1 2D model drawing of Model 1 (existing) [5]", + "texts": [], + "surrounding_texts": [ + "See Figs. 82.1, 82.2 and 82.3 and Tables 81.1 and 82.2. 82.3.3 Software Used 1. Modelling\u2014SOLIDWORKS 2016 2. Analysis\u2014ANSYS WORKBENCH R16. 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 975 976 E. Madhusudhan Raju et al." + ] + }, + { + "image_filename": "designv11_80_0002253_1350650120908116-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002253_1350650120908116-Figure12-1.png", + "caption": "Figure 12. Velocity vector diagram of the gear pump.", + "texts": [ + " The larger pump outlet pressure leads to a smaller export flow due to more serious leaks in the pump, so the trends of lines in picture conform to the actual test condition. Moreover, there exist errors of 3.07% and 5.87% under different pump inlet pressures. The local flow-rate offsets slightly because there is no escape from the occurrence of error due to the actual measurement. Standard k\u2013\" model and initial conditions of the gear pump should be set before the start of simulation. The pump rotational speed is n\u00bc 2400 r/min, and the input and output pressures are set as 0.07MPa and 2\u20139MPa. In Figure 12, the vector velocities and streamlines vorticity through gear pump are shown for one instant during the gear meshing cycle. The area investigated in picture is an axial plane cross pump flow domain. Note that the color of vectors represents the value of the flow speed. Through the zoom in narrow area, the flow speed in correspondence of the connection between the narrow areas is very large. It is known that the trapped-oil phenomenon easily occurs at the position of narrow area, so a detailed investigation should be done here. The position of monitoring point is also shown in Figure 12. Figure 13 shows the pressure distribution in pump with a rotation speed of 2400 r/min and different outlet pressures. The pressure gradually increases from the inlet to the outlet with the occurrence of partial high pressure in the trapped-oil area. The basic distribution law under different outlet pressure conditions is similar, but the pressure in the outlet and trapped-oil area becomes larger with larger output pressure. The pressure caused by oil trap is further simulated at the monitoring point in Figure 14" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003635_052061-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003635_052061-Figure5-1.png", + "caption": "Figure 5. Graphic representation of the blade bearing model", + "texts": [ + " Used blade coordinate system [12] To calculate dynamic rolling element loads the referenced double rowed four-point ball bearing was implemented in SIMPACK. The model consists of an outer ring rigidly connected to the inertial coordinate system with 0 DOF and an inner ring with 5 DOF coupled to the inertial system to prevent rotation around its middle axis. Both rings are connected by force elements equally distributed along the raceway to substitute each rolling element. A graphic representation of the blade bearing MBS model and the rolling element load distribution is displayed in Figure 5. The Science of Making Torque from Wind (TORQUE 2020) Journal of Physics: Conference Series 1618 (2020) 052061 IOP Publishing doi:10.1088/1742-6596/1618/5/052061 The dynamic load spectra acting on the inner bearing ring had been implemented by input functions. The dynamic rolling element loads during oscillation are calculated by interpolation taking the instantaneous position of each rolling element into account. Therefore, a cascade has been set up between the blade bearing model and the dynamic code", + " Therefore, dynamic accelerations due to centrifugal and inertial effects are neglected. Figure 6 shows the geometry of the rolling element-raceway interaction used in this approach, generally described by Gupta in \u201cAdvanced Dynamics of Rolling Elements\u201d [4]. To determine the contact vectors the dynamic model described by Gupta was adjusted to the shapes of the raceways of four-point contact ball bearings. Furthermore, it has been assumed that the inertial and the raceway coordinate systems have the same position. Due to the overlapping load zones of blade bearings (Figure 5), the resulting rolling element speed is calculated by vector addition of both diagonal The Science of Making Torque from Wind (TORQUE 2020) Journal of Physics: Conference Series 1618 (2020) 052061 IOP Publishing doi:10.1088/1742-6596/1618/5/052061 speed profiles defined by their contact angles. The circumferential speed of the linear profiles is described by tracing the contact angles for each raceway according to equation 1. \u20d7 , \u2219 ! \u2219 \"# \u2219 $\u20d7# , $\u20d7# , % $\u20d7& , (1) Thereby, $\u20d7# , $\u20d7& are the contact radii towards the bearing center axis, \"# is the rotational speed of the inner ring and is the rolling diameter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure5-1.png", + "caption": "Figure 5 Brake disc stress cloud map", + "texts": [ + " Defining constraints, apply full constraints on the center hole surface of the brake disc, apply X constraints on the inner and outer end surfaces of the disc, and apply Y and Z constraints on the friction surface of the brake disc and the friction lining. Its structure is shown in the figure below (Figure 4): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 The simulation calculation and analysis of the structural strength of the brake disc, and the results of stress solution and modal analysis deformation cloud diagram are shown in the following figure (Figure 5-6): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 From the strength calculation and cloud diagram of the brake disc structure, it can be seen that the maximum stress of the brake disc during the braking process is 52MPa, and the yield strength of the brake disc is 250MPa, so the brake disc fully meets the strength requirements. It can be seen from the deformation cloud diagram that the maximum deformation of the brake disc during braking is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003185_012018-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003185_012018-Figure2-1.png", + "caption": "Figure 2. Four-spring harvester with two spaced magnets.", + "texts": [], + "surrounding_texts": [ + "Sources, converting mechanical energy into electrical, such as the electromagnetic harvesters, are increasingly replacing the battery-type low power consuming power-supplying electronic devices [1, 2]. There are usually two types of magnet motion in relation to the coil in these harvesters. In the first case, the magnet moves in parallel to the coil [3] and in the second one - vertically with respect to it, often being located in its air gap [4]. Three different structures of four-spring electromagnetic harvesters are considered here, with two different masses, represented by steel plates with 2 or 4 magnets each and two different coils as well. The harvesters have a fixed coil attached to them, in parallel to which the rare earth magnets move. The aim of the present work is to study the effect of the construction parameters (such as the weight of the concentrated mass, the number of turns, the influence of the coil area, the volume of the magnets, and the size of the air gap) on the output electrical parameters of the studied harvesters. 2. Exposition TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 When applying a sinusoidally varying force with a resonant frequency, the mechanical system: \u201cmass (plate with permanent magnets) \u2013 springs\u201d begins to oscillate in parallel with the coil. Thus the magnetic flux passes differently through the coil and creates an alternating magnetic field in a different way, thus inducing alternating electromotive force. The studied harvesters are nonlinear mechanical oscillating systems. Their simulations made by ANSYS R19.1 take into account the fixture, the gravity effect of the plates with permanent magnets and the mechanical characteristics of the used springs. The horizontal deviation x of the mechanical system \u201cmass (plate with permanent magnets) \u2013 springs\u201d was obtained while modeling the four-spring electromagnetic harvesters, Figure 4. The magnetic field distribution of the three electromagnetic harvesters was obtained by means of FEMM 4.2. Figure 5 shows the magnetic field distribution for the third four-spring harvester with two spaced magnets in two places at zero horizontal deflection, and Figure 7 presents the distribution at maximum deviation. Figure 6 and figure 8 illustrate the magnetic flux density changes along the length of the harvester coil at zero and maximum horizontal deviation. From Figure 6 it can be seen that at zero horizontal deviation the normal magnetic flux density is zero, and at maximum deviation the maximum magnetic flux density Bmax is obtained. TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 The horizontal deviation at a sinusoidally changing force with a resonant angular frequency \u03c9 is also sinusoidal ( ) sinmx t X t\u03c9= (1) The horizontal oscillatory speed is equal to ( ) ( ) sin 2m d x t v t X t d t \u03c0\u03c9 \u03c9 = = + (2) The magnetic flux can be expressed by the change in the cross section of the coil \u0410(t), through which the maximum magnetic flux density Bmax passes ( ) ( ) max dA t \u0424 t B dt = (3) The change in the cross section \u0410(t) over time equals the horizontal deviation variation in the time x(t) along the diameter D\u0441 of the corresponding coil. ( ) ( ) \u0441 dA t D x t dt = (4) From (3) and (4), the magnetic flux through the coil is obtained ( ) ( )max \u0441 \u0424 t B D x t= (5) The induced electromotive force in the coil of the electromagnetic harvester is ( ) ( )d\u0424 t e t N dt = \u2212 (6) From (5) and (6) for the induced electromotive force for no-load mode, it is obtained ( ) ( ) max \u0441 d x t e t N B D dt = \u2212 (7) The amplitude of the induced electromotive force in the coil is a function of the amplitude of the oscillatory speed \u03c9Xm m max \u0441 m\u0415 N B D X\u03c9= (8) The angular frequency of the forced oscillations can be expressed by the frequency f 2 f\u03c9 \u03c0= (9) TechSys 2020 IOP Conf. Series: Materials Science and Engineering 878 (2020) 012018 IOP Publishing doi:10.1088/1757-899X/878/1/012018 From (8) and (9) for the amplitude of the induced electromotive force in the coil at a resonance it is obtained active load. In it, Rc and Lc denote the active resistance and the inductance of the electromagnetic harvester coil, \u0435c(t) - the induced electromotive force, and RL is the active load resistance. UL indicates the rectified voltage over the load resistance. The active power, in DC mode, is calculated using the amplitude of the induced no-load electromotive force in the coil at resonance, the voltage on the germanium diode UD and the parameters of the equivalent circuit (14). In the resulting expression, b denotes the attenuation coefficient, which is determined by the logarithmic attenuation decrement \u03b4 [5], where T is the oscillation periodic time and \u0415m(t) is the amplitude of the measured electromotive force 2b m\u03b4= , ( ) ( ) 1 ln m m E t T E t T \u03b4 = + (13)" + ] + }, + { + "image_filename": "designv11_80_0000542_978-3-030-20131-9_261-Figure98-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000542_978-3-030-20131-9_261-Figure98-1.png", + "caption": "Fig. 98. Bead string state diagram of touching the movable travel switch.", + "texts": [], + "surrounding_texts": [ + "In chapter 2, three methods of automatic knitting of bead mat are put forward. Among them, Warp and Weft Automatic Weaving Method is the most simple, but the cost is higher and its stability is poor. The lock stitch sewing weaving method is rea lly good, but there exists three difficulties applied in the machine: The cross-sectional size of the hook must be smaller than the size of the bead hole as the hook needle needs to completely pass through the bead, which makes it difficult to hook the string smoothly. It is difficult to guarantee the parallelism between the cross-section of the loop ring and the end of the bead of the first row in steps 4-6. When weaving a larger size beaded cooling pad, threading the string into the longer aligned transverse bead holes becomes hard. So compared with the other two methods, single-line straight-through method, which has good stability and high knitting efficiency and is easy to be realized on the machine, is the focus of the discussion below. Based on this method, an automatic weaving device capable of weaving a beaded cool pad is proposed and designed. 4.1 Feeding Device Design Before weaving the beaded cooling pad, since all the beading arrangements are disordered, it is necessary to design a device that puts the beads into the weaving state in an orderly manner, which is called feeding device. Referring to the feeding mechanism of the firecracker weaving [14-16], Fig.14 shows the schematic diagram of the designed bead feeding mechanism device. Before the device runs, all the beads are placed in the hopper and the two guiding wheels. A small number of longitudinally are placed in a horizontal arrangement. When the two guide wheels rotate in opposite directions, the beads in the hopper will be putted into the guide groove and conveyed to the front of the beading device in an orderly manner. In order to avoid a rigid collision between the feeding device and the ball transported device, the end of the guiding groove is made by a material with better elasticity. 4.2 Design and Working Principle of the Beaded Pad Weaving Device Fig. 15 demonstrated the beaded pad weaving device. Since the figure is only for explaining the movement process of the beaded pad weaving device, the feeding device is not shown in the figure. And there are 7 motors in this device. Control motor A controls the movement of the threading device. Control motor B controls the movement of the movable line of the downlink line. Control motor C controls the movement of the movable stroke switch. Control motor D controls the rotation of the output port and the braided port. Linear motor E S. Ouyang et al.2544 drives the up-line feed ball push block movement. Linear motor F pushes the braided beaded cool pad unit into the braided port. Linear motor G drives the linear motion of the downlink line feed bead block. Before the device is operated, the downlink threading is first performed. After the downlink threading is finished, the string is installed into the beaded pad weaving device. The downlink line with heavy beads at the end is wrapped around the fixed pulley mounted on the frame. Then it passes through the downlink line movable seat, the through hole, the downlink end sleeve, the bead and the braided port successively, to reach the uplink line. And the uplink line is directly connected to the needle of the threading device, the first string of beads are moved to the corresponding position on the weaving port, as shown in Fig.16. After the string installation is completed, the motor A is manually controlled to make the driving roller be located between the two trapezoidal blocks on the movable seat rail of the threading device. The manually controlled linear motor E is to drive the uplink line to send th e beads push block moves, which is external bead conveyed from the feeding device. It causes the holes axis of the bead to coincide with the needle axis of the bead threading device and the up-line bead push block is in the beading state. The specific work ing process of the device is as follows: Method Research and Mechanism Design of Automatic Weaving\u2026 2545 Step 1: Under the drive of the control motor A, the bead threading device moves to the right. When the driving roller moves in the second trapezoidal block on the movable seat rail, the location clamping position of the needle will be changed. As a result, the uplink line smoothly penetrates an external bead provided by the uplink line bead transported device, as shown in Fig.17. Step 2: The bead threading device continues to move to the right. When the movable seat contacts the movable travel switch, the uplink line is just tightened. At this time, some of the motor operation will change as follows: Control motor A reversed means the bead threading device starts to move to the left. Controlling motor B rotated forward means the downlink line movable seat moves to the right for a suitable distance, providing two beads re quired for the next unit downlink line weaving. After that, controlling motor B stops. Linear motor F runs, the weaving beads are pushed into the weaving port and the output port, then moves back to the initial position. Linear motor E reversely drives, the pushing block of uplink line feeding bead returns to the initial position, and is on out feeding condition, as shown in Fig.18. S. Ouyang et al.2546 Step 3: Under the driving of the control motor A, the bead threading device moves to the left. When the threading device contacts the fixed stroke switch, some of the motor operation will change as follows: Linear motor G forward drives, the pushing block of the downlink line feeding bead pushes a shared bead to make the hole axis of the shared bead coincide with the needle axis. Control motor A rotated forward means the bead threading device starts to move to the right. Linear motor E drives forward, the push block of the uplink line feeding bead is in the feeding state. Control motor C rotated forward, the movable travel switch moves to the left for a suitable distance exactly equal to the length of the string required to weave every bead pad unit, as illustrated in Fig. 19. Step 4: When the pushing block of the downlink line in the feeding state, the control motor D is drive to rotate the output port and the braided port counterclockwise by 180\u00b0. Step 5: Under the driving of the control motor A, the bead threading device moves to the right. When the driving roller moves in the first trapezoidal block on the movable seat rail, the location and clamping position of needle will be changed to make the uplink line successfully penetrate into a shared bead provided by the downlink line feeding device. Step 6: The linear motor E is reversely driven to make the pushing block of the downlink feeding bead out of the feeding state, as shown in Fig. 20. Method Research and Mechanism Design of Automatic Weaving\u2026 2547 Step 7: Repeat the actions from steps 1 to 6 until the end of the weaving task. Step 8: When the weaving process is finished, each motor is controlled by software programming to bring the device into an initial state." + ] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure5.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure5.6-1.png", + "caption": "Fig. 5.6 a Crash absorber box of the electric car prototype developed in the European project \u2018Evolution\u2019. It is made of rectangular aluminium alloy profiles filled with Al-foam. b CAD design of the body in white (Garcia-Moreno 2016)", + "texts": [ + " New trends and developments in the automotive industry, especially in the electric car segment, increase the demand for new concepts and materials for lightweight construction. Furthermore, new car designs are necessary due to the rearrangement of components, making it possible to consider cellular materials from the beginning. Passenger safety is another important factor, where a light, as compact as possible but very effective crash protection system is needed especially because the available crash space is reduced due to the absence of the traditional front engine. An example is shown in Fig. 5.6. Metal foam parts developed by the Technical University Berlin and Pohltec Metalfoam are foreseen in the prototype of an ultra-light electric vehicle recently developed in the European project \u2018Evolution\u2019 by a number of companies includingCidaut (Valladolid, Spain), Pininfarina (Cambiano, Italy) andPohltecmetal foam (Cologne, Germany). Railway industry is an important factor in future mobility concepts. Promising prototypes have evolved in the past years as possible future serial application. AFS foam panels delivered by the IWU (Chemnitz, Germany) have been used in the floor of a wagon of the metro in Peking in continuous operation without issues since 2008" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000236_0954407019838415-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000236_0954407019838415-Figure1-1.png", + "caption": "Figure 1. Schematic diagram of deformed crankshaft journal in the bearing hole.", + "texts": [ + " The lubrication equation21 used in the lubrication analysis of engine main bearings with rough surface considering the axial movement of deformed crankshaft can be described as follows \u2202 \u2202u fxh 3 \u2202p \u2202u +R2 \u2202 \u2202y fyh 3 \u2202p \u2202y =6hUR \u2202hT \u2202u +6hVR2 \u2202hT \u2202y +6shUR \u2202fs \u2202u +6shVR2 \u2202fs \u2202y +12R2h \u2202hT \u2202t \u00f01\u00de where u is the bearing circumferential coordinate; h is the oil film thickness; p is the oil film pressure; y is the bearing axial coordinate; U is the tangential velocity of journal surface, U=pRn=30; n is the crankshaft speed; h is the lubricating oil viscosity; R is the bearing radius; V is the axial velocity of crankshaft; s is the surface comprehensive roughness, s= ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi s2 1 +s2 2 q ; s1 and s2 are the journal and bearing surface roughness, respectively; fs is the shear flow factor; fx and fy are the pressure flow factors; and hT is the gap between the rough surfaces21,22 hT = h 2 1+ erf hffiffiffi 2 p s + sffiffiffi 2 p p exp h2 2s2 where erf () is the error function. As shown in Figure 1, when the crankshaft under load deformation causes the shaft journal to tilt in the bearing hole, the movement of deformed crankshaft along the bearing axis directly affects the distribution of oil film thickness in axial section of bearing, and further changes the distribution in the radial section. The oil film thickness equation considering the axial movement of deformed crankshaft can be written as h= c+ e0 cos u c\u00f0 \u00de+ tang y+V Dt L 2 cos u a c\u00f0 \u00de \u00f02\u00de where e0 is the eccentricity on the central section of bearing, a is the angle between the projection of the center line of the rear end of the crankshaft journal and the eccentricity vector, c is the angle of displacement of the central section of the bearing, c is the bearing clearance, g is the inclination angle of crankshaft in bearing hole, Dt is the time required for the unit angle of the crankshaft rotation, and L is the bearing width", + " 3 main bearing is increased by 5.87%. This is mainly due to the tilt angle and azimuth are significantly different in different main bearings, in the same axial displacement and different misaligned condition coupling effect, so that the radial displacement components of the crankshaft journal center caused by the axial movement of deformed crankshaft are quite different. The minimum oil film thickness decreases when the journal tilts to the left and increases when the journal tilts to the right when combined with Figure 1. Therefore, the influence of crankshaft axial movement on the minimum oil film thickness of different main bearings is directly related to the tilt angle, azimuth and axial movement direction and displacement of deformed crankshaft journal. The hmin of No. 1 and No. 3 main bearings are 0.9905 and 0.4696mm, respectively, when the axial movement of deformed crankshaft is taken into account without considering the influence of the surface roughness. When the axial movement of deformed crankshaft and the surface roughness are taken into account, the hmin of No" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002906_pen.25409-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002906_pen.25409-Figure8-1.png", + "caption": "FIGURE 8 Predicted parison shape at Weissenberg number Wi = 40 and for a die exit cross-sectional aspect ratio \u03c7 = 12.5, using a viscoelastic PTT model, A, perspective view of the extruded sheet with sag effect, B, cross-sectional shape of the extruded sheet at a distance of 60 mm from the die exit showing thickness and width swelling", + "texts": [ + "[35] Let us first consider the extrusion from a rectangular die exit cross section with an aspect ratio \u03c7 of 12.5 (that is, thickness h = 8 mm and width L = 100 mm). The inlet velocity profile is given by Equation (39), with Vmax = 5 mm/s, that corresponds to a Weissenberg number of about Wi = 40. The gravity is also included in the simulation. We show in Figure 7 the predicted parison shape at various times during extrusion. Throughout the extrusion, the sag due to the parison weight becomes prominent (time = 333 seconds), and a swelling reduction is observed at mid-length of the extrudate. In Figure 8, we show the extrudate swell in the cross section, at z = 60 mm from the die exit. The thickness swelling ratio is 1.45, while the width swelling ratio is 1.34. It is worth to mention that these swelling ratio are not uniform along the extrudate due to the sag. The distribution of the first normal stress difference N1 = \u03c4zz \u2212 \u03c4yy along the central line, in the flow axis, is presented in Figure 9A (the flow is along the z-axis while the shear is oriented with the y-axis). It starts from zero in the die, increases to reach a maximum just before the exit and then decreases to become negative just before the die exit" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002337_eiconrus49466.2020.9039215-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002337_eiconrus49466.2020.9039215-Figure7-1.png", + "caption": "Fig. 7 The stator with distributed winding", + "texts": [ + " In this paper, we propose to consider an electric machine with a distributed four-layer winding with an open groove as a scooter wheel engine. Such a groove configuration greatly simplifies the stator winding procedure of an electric machine [5]. We apply the numerical method developed here to simulate the magnetic field of permanent magnet electric machines. As an example, consider the engines used in the motor-wheel drive of electric scooters. One of these motors has a concentrated winding (Fig. 6). The second engine has a distributed winding developed by the authors (Fig. 7). Both engines have the same dimensions and the same power, but they differ greatly in electrical parameters. Table 1 shows the main parameters of the compared electrical machines. TABLE. COMPARISON OF ENGINE PARAMETERS Parameter Type of winding concentrated distributed Rated power, Watt 1200 1200 Rated voltage (DC), V 60 30 Rated current, A 20 40 Rated torque, N \u00d7 m 10 10 Max torque, N \u00d7 m 15,7 15,8 Rated speed rp/m 1000 1000 The number of pole pairs 15 8 Rated current frequency, Hz 250 133 Active phase resistance, mOhm 67 17 Inductance, \u03bcH 270 19,6 The mass of magnets, g 335 305 From the above data it is seen that the distributeddeveloped motor developed by the authors has advantages over the currently widely used concentrated-winding motor model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001429_j.mechmachtheory.2019.103674-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001429_j.mechmachtheory.2019.103674-Figure3-1.png", + "caption": "Fig. 3. Haptic Feedback System RePlaLink .", + "texts": [ + " Its kinematic structure is a reconfigurable planar linkage, designed for the haptic simulation of mechanisms. For the universal haptic simulation of most hand-actuated mechanisms, three DOF in the plane are required, as most commercial hand-actuated linkages are planar. Because HFS have to be able to render high stiffness orthogonal to the coupler curve, it is advantageous to use HFS with only three DOF in plane. The kinematics can be adjusted in height and orientation to adapt the workspace to the application, see Fig. 3 (a). The hybrid structure consists of a parallel five-bar mechanisms for positioning and an additional sixth serial link for rotating the handle. A programmable logic controller (PLC) is used for control and simulation with a cycle time of 1.2 ms. It has only a moderate computation power with an Intel Atom Z520 and 1.3 GHz motivating an efficient implementation of all required models and calculations. The industrial drive system includes three permanently excited servo drives and appropriate servo inverters", + " Humans are capable of exerting high forces around 100 N and it is expected that mechanisms partially require similar forces [35] . For that reason motors with the following properties are chosen: The nominal torques are 77 Nm at a rated speed of 280 rpm for the bigger servos and 10 Nm at a rated speed of 94 rpm for the small servo. The six-axis force and torque sensor attached between the handle and the rotating link measure forces exerted by the user. An overview of the drive system, the kinematic structure and the most relevant dimensions is shown in Fig. 3 (b). The direct kinematics can be solved by solving the dyad ACB with the half-tangent-method ( [36] , p. 412). Defining the minimal coordinates q as the drive angles (1) , the direct kinematics (3) are the mapping from q to the pose (position in x and y as well as orientation) x H (2) of the handle H and the inverse kinematics vice versa (4) . q = [ \u03d5 1 \u03d5 2 \u03d5 3 ]T (1) x H = [ x H y H \u03d5 H ]T (2) x H = DK ( q ) (3) q = IK ( x H ) (4) Formulating the closing condition for the velocities leads to the Jacobian J and the overall Jacobians for all centres of gravity (cog) for the translations J and the rotations J " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000821_ccaa.2018.8777578-Figure19-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000821_ccaa.2018.8777578-Figure19-1.png", + "caption": "Fig. 19: Top view of the arena 4.", + "texts": [ + "The arena consists of a circle placed a height of 548.64cm from the ground. This arena was made to test the altitude stability. TABLE III: OBSERVATION OF THE ARENA 3. Trial Time(S) 1 15 2 17 3 12 4 15 5 14 Table III shows the number of trials performed and time taken to complete the task. This test is performed to find quad\u2019s altitude stability, as the quad need to be in the level of the circle to pass through it. The maximum time where the quad had a stable altitude was for 5 seconds. The best flight obtained was in 12 seconds. Arena 4: Fig. 19 shows the top view of the arena 4 and Fig. 20 shows the practical implementation of the arena. The arena consists of two hurdles: one is circle and the other is a square. The circle if is of diameter 100cm and the square with side 100cm.The circle and the square are held at a distance of 180 cm. The hurdles were set at a height of 6 feet from the ground. Circle was the first hurdle which was tied to the roof bars of the corridors and the square hurdle was suspended from the roof bars as shown in the Fig. 19. This arena was setup to test quad\u2019s motion error while travelling in the straight path. Table IV shows the number of trials performed and time taken for completing the task. This arena was setup in such a way it will test quads stability and to find any errors in its straight line motion. The best flight in this arena was in 33sec. Arena 5: Fig. 21 shows the top view of the arena 5 and Fig. 22 shows the practical implementation of the arena. This arena is mixed one, through which we were able to test quads stability, motion errors" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003452_0954405420951101-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003452_0954405420951101-Figure7-1.png", + "caption": "Figure 7. Calculation model of maximum root stress recommended ISO.", + "texts": [ + " The contact ratio can be calculated as: e= z1 tanaa1 tana1\u00f0 \u00de+ z2( tanaa2 tana2)\u00bd =2p \u00f012\u00de The radius of the switching point between the single teeth-meshing area and the double teeth-meshing area can be calculated by: rj1= ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r2b1+ ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi r1 + ha1m\u00f0 \u00de2 r2b1 q (e 1)pmcosa1 2s \u00f013\u00de The pressure angle on this point can be obtained by: aj1 = arccos (rb1=rj1) \u00f014\u00de The most common mathematic model to calculate the maximum tooth root stress is 30 incline tangent method recommended by ISO,15 which is shown in Figure 7. The inscribed profiles are oblique line 30 from the vertical line. Critical section location is determined as the tangent point of the gear profile and inscribed profile. The distance between two tangent point is defined as Sf. The vertical distance between the stress point on the symmetry axis and tangent point is noted as hf. The force loaded on the switching point between single tooth engagement and double tooth engagement is noted as F. The maximum bending moment can be calculated as: T=F hf cos (a+b) \u00f015\u00de where b can be obtained by: b= tan aj tan(a) p=2z1 \u00f016\u00de The tooth width is noted as l and the bending section modulus of rectangle can be expressed as: W= l Sf 2=6 \u00f017\u00de The bending stress can be evaluated by: sf =T=W= F hf cosaj1 l Sf 2=6 \u00f018\u00de By establishing a coordinate system at the center of the gear, hf can be obtained by calculating other parameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002806_012079-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002806_012079-Figure2-1.png", + "caption": "Figure 2. the position of a wheel on brake roller tester [2]", + "texts": [ + " BIS-ASE 2019 Journal of Physics: Conference Series 1517 (2020) 012079 IOP Publishing doi:10.1088/1742-6596/1517/1/012079 The mechanical drawing of roller brake tester is shown in Figure 1. It consists of main elements, rotating directions, forces and dimensions of the roller brake tester. The information presented in Figure 1 is very essential for understanding the working principle of brake tester system and for determination of the braking force. The appropriate position of the vehicle wheels placed on the brake roller tester is shown in Figure 2. When the brake roller tester runs, both rollers rotate at a constant speed (\u03c9W1 = \u03c9W2). These rollers are driven by a constant torque electromotor. When braking force is applied, the wheel decelerates and a torque MW is developed. The torque lever which is connected to the stator part of the electromotor deflects in the opposite direction from the rotations of rollers, electromotor rotor (\u03c9R), and vehicle wheel (\u03c9W). The torque resulted from the vehicle wheel (MW) is transmitted through roller sprockets and chain, electromotor rotor and finally to electromotor stator (lever)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002331_humanoids43949.2019.9034995-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002331_humanoids43949.2019.9034995-Figure12-1.png", + "caption": "Fig. 12. Representation of the airbag as a nonlinear spring-damper system.", + "texts": [ + " By computing this transition, we can provide the new initial conditions for the motion calculation with the following pendulum. Traditionally, the impact is modeled as an inelastic collision, and the transition equations can be analytically derived [14]. However, in our case the effect of the airbags needs to be included. The airbag can be modelled as a spring-damper system. The elements, however, do not behave linearly; the spring stiffness k(x) and the damping coefficient c(x) depend on the compression x of the airbag (see Fig. 12). The equivalent spring coefficient k(x) is obtained from the derivative of the impact force F (x) with respect to x. k(x) = dF (x) dx (1) The damping coefficient is computed from the stiffness k(x), the damping ratio h, and the natural frequency \u03c90 as c(x) = k(x) 2h \u03c90 . (2) The impact force F (x) equals the product of the internal airbag pressure p(x) and the contact area A(x). Note that the contact area is also a function of the deformation x, which in turn is a function of time, x(t). To simplify the notation, the explicit dependency on time will be dropped hereafter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure90.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure90.4-1.png", + "caption": "Fig. 90.4 Load and boundary conditions assigned on liner", + "texts": [ + " Hence, weld bead + HAZ is considered as a single region and measured to be 15 mm in length. In the present analysis, weld material properties are assigned for a length of 15 mm as shown in Fig. 90.3b in the pole region and Fig. 90.3c in the equator region. The symmetric boundary condition is applied on side faces of liner by restricting Z-axis movement (U3 = 0). Also, bottom face of liner is restricted by Y-axis movement (U2 = 0) and inlet face movement is restricted by X-axis movement (U1 = 0), respectively as shown in Fig. 90.4. The internal pressure load is applied on the interior face of liner (refer Fig. 90.4). Liner burst test is conducted to determine its pressure load contribution and check the failure location due to applied load. The applied burst pressure load on the liner is determined using the Eq. 90.1. P \u00bc 2 r t R P \u00bc 2 290 1:5 210 \u00bc 4:15MPa \u00f090:1\u00de where r \u00bc Ultimate tensile strength of AA 6061-T6, t = Minimum thickness of liner and R = Radius of liner. Dynamic explicit analysis is conducted with non-linear geometric effects. For prediction of failure, Johnson\u2013Cook (JC) damage material model is utilized" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002315_012062-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002315_012062-Figure2-1.png", + "caption": "Figure 2. Tangential creep effects according to [5].", + "texts": [ + " As the described sequence occurs with every overroll, roller-induced creep increases with rotational speed and time. In addition to roller-induced creep, gear creep can occur if the bearing housing is a gear. The radial and tangential load situation in each gear tooth engagemet and the resulting tooth bending lead to the formation of an additional gap in the area of each teeth enagement. When set in rotation, this gap is set into motion as well. Other than roller-induced creep, gear creep drives the bearing ring against the rotational direction of the rolling element set. Figure 2 shows the rotational direction of the bearing ring (red dot) relative to the gear wheel (blue dot) for roller-induced creep (middle) and gear creep (right). The rotational direction of the rolling element set (black dot) is the same in all pictograms. NAWEA WindTech 2019 Journal of Physics: Conference Series 1452 (2020) 012062 IOP Publishing doi:10.1088/1742-6596/1452/1/012062 In [5] it was found out that gear creep already occurs at lower loads than roller-induced creep. However, the speed of roller-induced creep is higher" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure61.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure61.4-1.png", + "caption": "Fig. 61.4 Surface of die and chamber (a), spiral face of die, (b) and cavities of chamber (c)", + "texts": [ + " Then after the sample bar extruded, the die was removed by engaging the clash downward to change feed direction. 61 Fabrication of Hexagonal Bar from Aluminum Alloy AA6063 \u2026 739 The scroll-faced die and one side cavity chamber are the basic equipment in friction stir back extrusion (FSBE). The function of this die and chamber is to consolidate, stir, and extrude aluminum scrap. To resist wear occurrence in the process of stirring and hexagonal-shaped sample bar production, it should be heat treated. The color looks like black after heat treatment as Fig. 61.4. Figure 61.5 shows the samples fabricated at different rotational speed and their shapes at each rotational speed. For example, sample one is extruded at 630 rpm and 6.25 mm/min, sample two 500 rpm and 6.25 mm/min, etc. Figure 61.5 shows the possibilities of FSBE to fabricate hexagonal-shaped bar and other shapes from aluminum alloy scrap. Most of extrudes twisted and circular shapes. The reason of twist is, in FSBE die shape, machine power and extruded material properties affects the shape, length, and mechanical properties of extrudes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000293_012101-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000293_012101-Figure4-1.png", + "caption": "Figure 4. Total deformation: structural steel (a), alloy steel (b) and carbon steel (c).", + "texts": [ + " Carbon steel had the equivalent stress, maximum principal stress and maximum shear stress of 933.23, 408.74 and 482.44 MPa. Overall, among the other materials under study, alloy steel had a lower strength. A drop in stress indicates a higher strength [15]. In sum, structural steel is considered to have the highest strength among the three materials. IC2MAM 2018 IOP Conf. Series: Materials Science and Engineering 515 (2019) 012101 IOP Publishing doi:10.1088/1757-899X/515/1/012101 The simulation results of the total deformation of the three starter driven gears are shown in Figure 4. Maximum equivalent stress decreases as the deformation rate decreases [16]. This is an evident that alloy steel has the lowest deformation along with its stress. The crack occurring in each starter driven gear was examined based on the values of J-Integral and SIFS (K1). J-Integral value can be described as the strain energy release rate of a crack body per unit. Moreover, the SIFS (K1) value serves to determine the stress intensity factor (K) of a material with a certain geometry shape under elastic loading" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001688_auto-2019-0078-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001688_auto-2019-0078-Figure2-1.png", + "caption": "Figure 2: Illustration of vehicle model.", + "texts": [ + " The dynamics of each vehicle i \u2208 N is described by a simple model (bicycle model, e. g., [16]), i. e., \u2200t \u2208 \u211d\u22650, {{{ {{{ { x\u0307i(t) = vi(t) cos(\u03d1i(t)) y\u0307i(t) = vi(t) sin(\u03d1i(t)) \u03d1\u0307i(t) = vi(t) L tan(\u03b2i(t)) , (1) where xi(t) \u2208 \u211d\u22650 and yi(t) \u2208 \u211d\u22650 are longitudinal and lateral positions of the vehicle\u2019s center of mass, vi(t) is its velocity, \u03d1i(t) the yaw angle, and \u03b2i(t) the steering angle. Velocity vi(t) and steering angle \u03b2i(t) are considered to be the control inputs. L is the vehicles\u2019 length (for ease of notation, all vehicles are assumed to have the same length). Figure 2 illustrates themodel parameters. In what follows, pi(t) := [xi(t), yi(t)] denotes the position of vehicle i \u2208 N on the highway, with pi(t0) := [xi0 , yi0 ] its initial position. Each vehicle i \u2208 N has a desired velocity vdi \u2208 [vmin, vmax]. Since the controller will be implemented digitally (MPC), system (1) is sampled with a sampling time Ts \u2208 \u211d>0. We denote \u2200k \u2208 \u21150, xi(k) := xi(kTs). (2) (similarly yi, \u03d1i, vi, and \u03b2i). Let, \u2200i \u2208 N , \u2200k \u2208 \u21150, qi(k) be the lane in which vehicle i is at time instant k" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure22-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure22-1.png", + "caption": "Figure 22. Shear stress distribution from a plain bearing to a textured surface", + "texts": [], + "surrounding_texts": [ + "CoNCLUSIoN\nThis\ufeff research\ufeff presents\ufeff a\ufeff study\ufeff on\ufeff the\ufeff textured\ufeff surface\ufeff effect\ufeff in\ufeff the\ufeff hydrodynamic\ufeff plain\ufeff journal\ufeff bearings.\ufeffFor\ufeffimprovement\ufeffof\ufeffelastohydrodynamic\ufefflubrication\ufeffperformance\ufefffor\ufeffthe\ufeffboth\ufefftype\u2019s\ufeffjournal\ufeff bearings:\ufeffwith\ufefftextured\ufeffsurfaces\ufeffand\ufeffuntextured\ufeffsurface,\ufeffusing\ufeffnumerical\ufeffmodeling\ufeffimplement\ufefffinite\ufeff element\u2019s\ufeffmethod,\ufeffa\ufeffpara-\ufeffmetric\ufeffstudy\ufeffwas\ufeffperformed\ufeffto\ufeffinvestigate\ufeffthe\ufeffeffect\ufeffto\ufeffvarious\ufeffoperating\ufeff conditions\ufeffon\ufeffpressure,\ufeffdisplacement\ufeffand\ufeffshear\ufeffstress:\n\u2022\ufeff The\ufeffresults\ufeffwere\ufeffcompared\ufeffwith\ufeffthe\ufeffbearing\ufeffnot\ufefftextured\ufeffand\ufeffbearing\ufeffentirely\ufefftextured.\ufeff The\ufefffollowing\ufeffare\ufeff the\ufeffkey\ufeffconclusions\ufeffdrawn\ufefffrom\ufeffthe\ufeffstudy:\ufeffThe\ufeffpressure\ufeffdistribution\ufeff according\ufeffto\ufeffangular\ufeffposition\ufefffor\ufeffbearing\ufefftextured\ufeffand\ufeffnot\ufefftextured\ufefffor\ufeffradial\ufeffload\ufeffof\ufeff10\ufeff 000N\ufeffand\ufeffrotational\ufeffvelocity\ufeffof\ufeff9000rpm,\ufeffhas\ufeffthe\ufeffsame\ufeffpattern\ufefffor\ufeffthe\ufefftwo\ufeffcases\ufeffstudied.\ufeff The\ufeffdifference\ufeffis\ufeffestimated\ufeff75\ufeffpercent; \u2022\ufeff The\ufeffevolution\ufeffof\ufeffthe\ufeffdisplacement\ufeffalong\ufeffthe\ufeffangular\ufeffposition\ufeffof\ufeffthe\ufeffbearing\ufeffhas\ufeffthe\ufeffsame\ufeffpattern\ufeff for\ufeffthe\ufeffnon-textured\ufeffbearing\ufeffand\ufefftextured\ufeffbearing,\ufeffthe\ufeffdifference\ufeffis\ufeffestimated\ufeffby\ufeff42\ufeffpercent; \u2022\ufeff The\ufeffmaximum\ufeffdeformation\ufeffis\ufeffnoted\ufeffat\ufeffthe\ufefflevel\ufeffof\ufeffthe\ufefflower\ufeffgeneratrix,\ufeffas\ufeffwell\ufeffas\ufeffthe\ufeffdisplacement\ufeff variation\ufeffof\ufeffbearing\ufeffwith\ufeff the\ufeff rotational\ufeffvelocity\ufeffvariation\ufeff is\ufeff significant\ufeff for\ufeffa\ufeffbearing\ufeffwith\ufeffa\ufeff textured\ufeffsurface\ufeffcompared\ufeffto\ufeffthat\ufeffobtained\ufefffor\ufeffa\ufeffnon-textured\ufeffbearing; \u2022\ufeff Maximum\ufeffvalues\ufeffof\ufeffthe\ufeffshear\ufeffstress\ufeffare\ufeffnoted\ufeffat\ufeffthe\ufeffangular\ufeffposition\ufeff60\u00b0\ufeffand\ufeff230\u00b0,\ufeffthe\ufeffdifference\ufeff between\ufeffthe\ufeffflow\ufeffrate\ufeffof\ufeffa\ufeffnon-textured\ufeffbearing\ufeffand\ufeffa\ufefftextured\ufeffbearing\ufeffis\ufeff42\ufeffpercent.\nACKNoWLEdGMENT\nThis\ufeffresearch\ufeffreceived\ufeffno\ufeffspecific\ufeffgrant\ufefffrom\ufeffany\ufefffunding\ufeffagency\ufeffin\ufeffthe\ufeffpublic,\ufeffcommercial,\ufeffor\ufeffnotfor-profit\ufeffsectors." + ] + }, + { + "image_filename": "designv11_80_0002002_robio49542.2019.8961771-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002002_robio49542.2019.8961771-Figure3-1.png", + "caption": "Figure 3. The shape of the continuum robot is divided into three parts by posture sensors. There are Shape of Part-1, Shape of Part-2 and Shape of Part-3, respectively. Each part will use shape reconstruction method based on cubic B\u00e9zier curve.", + "texts": [ + " These two functions can be expressed as: i si i i-1 l L / n i i 1 l = B B B( ) B( ) n n where n represents the number of points used to estimate the length of a curve in Quadratic B\u00e9zier curve fitting, which is equal to half of the number of the connection disks, then the error estimation function f can be expressed as: n 2 i si i 1 f ( l l ) Two unknown parameters ( lAB and lCD) can be calculated by finding the optimal solution of the optimal objective function. The LM algorithm is used to optimize the two parameters. The shape of one segment of the continuum robot can be detected by the above method. As shown in Figure 3, the shape of the continuum robot is divided into three parts by posture sensors. According to the method of piecewise fitting curves, the explicit form of the shape reconstruction is: 3 2 2 3 C1 1 1 1 1 3 2 2 3 C2 1 2 2 2 3 2 2 3 C3 2 3 3 3 B (t ) (1 t ) A 3t(1 t ) B 3t (1 t )C t D B (t ) (1 t ) D 3t(1 t ) B 3t (1 t )C t D B (t ) (1 t ) D 3t(1 t ) B 3t (1 t )C t D In this way, the shape of continuum robot can be reconstructed in 3D space even if the end of the robot is subjected to external load" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001685_icems.2019.8921850-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001685_icems.2019.8921850-Figure5-1.png", + "caption": "Fig. 5. Distribution of magnetic line of force and zoomed in view. (a) Normal. (b) 15-turn ITSC fault.", + "texts": [ + " When the ITSC fault occurs, the symmetry of the curve is destroyed, the amplitudes of the air gap magnetic density increase and fluctuate obviously, as well as the harmonic content increases significantly. The fast Fourier transform of air gap magnetic flux density is applied in two conditions, and the result is shown in Fig. 4. The magnetic density of the healthy motor only contains the harmonic orders of 6k\u00b11, but the fault will increase the amplitudes of harmonics and generate additional harmonics such as 3rd, 4th, 6th, etc. Distribution of magnetic line of force pre/post fault is shown in Fig. 5. It can be found that there is an obvious difference in the distribution of the magnetic line of force before and after the fault, that is, the fault breaks the symmetry distribution of the magnetic field. Compared with the healthy motor, a higher saturation phenomenon occurs around the faulty slot which is caused by an additional magnetic field generated by the short circuit current. The magnetic flux density distribution of the motor pre/post fault under the rated load is shown in Fig. 6. The overall appearance of the magnetic field is symmetrically distributed in two poles in healthy condition" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002751_s00162-020-00532-0-Figure16-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002751_s00162-020-00532-0-Figure16-1.png", + "caption": "Fig. 16 The instantaneous wake structure of the projectile under the condition ofMa = 3 and \u03b1 = 6\u25e6", + "texts": [ + " Figure 14 shows the pressure distribution on the plane of xoy under the condition ofMa = 3 and \u03b1 = 6\u25e6. Figure 15 shows the instantaneous Mach number contours of two kinds of projectiles. As seen from Fig. 14, the overall flow field structures of the two projectiles are similar, but the pressure distribution in the circular area of the warhead site and the tail position of the projectile is slightly different. The pressure difference of wake field is easy to understand, because it changes from a primary expansion wave to two expansion waves (as shown in Fig. 16). The mechanism for the evolution of the high-pressure area on the warhead is relatively complex. In Fig. 15, the velocity near the windward side increases due to the change in the boattail structure. As a result, the stagnation zone on the windward side of the warhead extends backward, leading the high-pressure aero to expand. Therefore, the drag and lift coefficients of the projectile increase with the addition of boattail structure. From Figs. 14 and 15, the Mach number and pressure contour on the leeward side are basically the same, and the flow field around leeward side is not affected much" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure16.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure16.6-1.png", + "caption": "Fig. 16.6 Modelled tensile test specimens with mesh, a BM and b WM", + "texts": [ + " One end of the specimen is ENCASTERED (U1 = U2 = U3 = UR1 = UR2 = UR3 = 0), where movement is restricted in all degrees of freedom, while another end is allowed to move in the X-direction for BM and WM (refer Fig. 16.5a, b). 16 Activated TIG Welding of AISI 321 Austenitic Stainless Steel \u2026 185 The meshing of the model is conducted by employing fine mesh at the gauge length section and coarse mesh at the gripping section (to reduce the computational time) as results are expected only at gauge length section. The parameters used for meshing in FE analysis are shown in Table 16.9. Further, the meshed tensile test model of BM and WM is described in Fig. 16.6a, b separately. Figure 16.7a shows the experimental tensile test depicting the fracture behaviour of the BM. Similarly, a tensile test conducted using the FE model is shown in Fig. 16.7b with a subsequent fracture. A good match can be observed in fracture at the centre of the specimen in both the analyses. Also, the fracture behaviour of the BM and WM samples was observed with the ductile mode of fracture and found similar with FE analysis as shown in Figs. 16.8 and 16.10. Likewise, for WM, the fracture region obtained after performing experimental and FE analysis is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001793_tmag.2019.2950181-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001793_tmag.2019.2950181-Figure1-1.png", + "caption": "Fig. 1. Approximation of one symmetry sector of a toroidal inductor by two axial (in green) and two radial (in blue) slices.", + "texts": [ + " Section II covers the theoretical aspects of the MASM. It also includes details about the geometry and supporting technical aspects for the simulations. Along with the necessary results, Section III provides in-depth comparative analysis followed by conclusions in Section IV. We widely use the definitions explained in [22] and [23]. 0018-9464 \u00a9 2019 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. Fig. 1 shows the considered toroidal inductor as well as one of its symmetry sectors. Both MASM and 3-D FEM are built for the symmetry sector. A linear ferrite core with relative permeability \u03bcr = 3000 is considered. The inner and outer radii of the core are designated by rm and rout, respectively. The height of the core is h. A copper conductor with 0.6 mm diameter is used. The inductor carries N = 50 turns equally distributed over the periphery as shown in Fig. 1. The symmetry sector covers an angle of \u03d5sym = 2\u03c0 /N . In Fig. 1, the symmetry sector is sliced with four planes, two axial ones in green (z = constant) and two radial ones in blue (r = constant), so that the current-carrying conductors are approximately perpendicular to each slice. The axial slices allow accounting for the magnetic core as well as the conductors at the inner and outer sides of the core. The radial slices allow accounting for the stray field on the top and bottom of the core. The core is only considered in the axial slices, not in the radial ones" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003327_speedam48782.2020.9161896-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003327_speedam48782.2020.9161896-Figure5-1.png", + "caption": "Fig. 5. Magnetic flux flow.", + "texts": [ + " Downloaded on October 17,2020 at 20:29:58 UTC from IEEE Xplore. Restrictions apply. Permalloy is an alloy of iron and nickel, and has high magnetic permeability. Furthermore, the saturation magnetic flux density of 50H800 used in the electromagnetic steel sheet is 1.82T, and permalloy is 0.8T. Utilizing the high magnetic permeability of permalloy and the difference of saturation magnetic flux density between permalloy and steel sheet, the high-speed range can be extended. Next, the magnetic flux flow in the low-speed and highspeed regions is described. Fig. 5 shows the flow of magnetic flux in each operating region. The high torque is needed in the low-speed range under heavy load, and therefore, it is necessary to much flow the magnetic flux from the magnet to the stator. Though the part of magnetic flux flows to permalloy, most of magnetic flux flows to the stator because the saturation magnetic flux density of permalloy is low and permalloy plays a role of flux burrier. Moreover, the difference in d-axis and q-axis inductance is larger than that of the variable leakage flux motor since the q-axis magnetic path is not blocked and the reluctance torque can be utilized" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001761_012021-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001761_012021-Figure1-1.png", + "caption": "Figure 1. The series of gear systems used in the hand tractor and one of the spur gears has experienced a fracture (indicated by red circle).", + "texts": [ + "1088/1742-6596/1351/1/012021 The chronology of the failure of the spur gear, which is used on the hand tractor, can be explained that on June 26, 2017 a tooth fracture occurred in the spur gear pair, where the tractor had been operating for about 40 hours prior to failure occurring. Therefore, this study aims to investigate and find the stress value that occurs in the gear that causes failure of the gears on a hand tractor using the finite element method. This study examined the spur gears of an incapacitated hand tractor after operating for only 40 hours, classified as premature failure. This research was conducted using the finite element analysis. The issue of this study is the fracture gear of gear transmission system used on a hand tractor as illustrated in Fig. 1. First of all, a visual inspection of the hand tractor spur gear is carried out to observe and identify the broken surface of the spur gear, as shown in Fig. 2. It needs to be identified is the characteristic of static or fatigue fractures that are usually marked by the beach mark on the surface of the fracture. URICSE Journal of Physics: Conference Series 1351 (2019) 012021 IOP Publishing doi:10.1088/1742-6596/1351/1/012021 Hereafter, chemical analysis was also carried out to determine the chemical composition of the gears so that it could be classified correctly into the proper standard groups, such as ASM or AISI" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003674_msm49833.2020.9201736-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003674_msm49833.2020.9201736-Figure2-1.png", + "caption": "Fig. 2. Kinematic chain of lower limb structure.", + "texts": [ + " Chapter VI discuss results of simulation experiments. Finally, a discussion is offered to summarise the paper. For the purpose of considered model an 8-DOF robot mechanism is proposed. Both ankle and knee joints has been modeled as 1-DOF mechanism. Movement of them is described in sagittal plane. The hip has been modelled as 2- DOF joint, describing movement in both sagittal and coronal plane. With view to simplifying kinematic chain, it is assumed that mass and length of connector between joints in the hip is equal to 0. Figure 2 depicts proposed kinematic chain structure. Base frame is set at the end of right foot, whereas end effector is set at the left ankle joint. Proposed kinematic chain is suitable for swing phase of gait, when the left limb is the swung one. Coordinate systems of the frames are assigned according to positive direction of revolution joints including Denavit \u2013 Hartenberg notation. During the experiment human gait was recorded to further identification of dynamic parameters, that will be used as input data to analyze proposed dynamic model of exoskeleton" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003937_ecce44975.2020.9235661-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003937_ecce44975.2020.9235661-Figure7-1.png", + "caption": "Fig. 7. 2D FEA IM cross section.", + "texts": [], + "surrounding_texts": [ + "The circuit parameters are not uniquely defined and depend on k. Thus, to match the physical conditions more closely, the static 3D FEA results are used to calculate the main inductance. The radial component of the magnetic flux density penetrating the rotor surface facing one pole of the machine is hereby defined as the main magnetic flux density. L\u00b5 is then calculated using the number of turns N : \u03a8\u00b5 = N \u222b ~Br \u00b7 d ~A (7) \u03a8\u00b5 Is,nl = L\u00b5 (8) The ratio k is then calculated using (6)." + ] + }, + { + "image_filename": "designv11_80_0003991_s40964-020-00153-3-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003991_s40964-020-00153-3-Figure5-1.png", + "caption": "Fig. 5 FEM simulations of the deformed female mould when subjected to water steam pressure at the process working temperature (the legend refers to the total deformation, in mm). Model with an equivalent material simulating the mechanical behaviour of the lattice (a) and without lattice structure in the conveying chamber (b)", + "texts": [ + " As already stated, the lattice structure also has to be designed to sustain the moulding loads at the proper processing temperature. The analyses have been conducted using a finite element model (FEM) by the ABAQUS software, considering two simplified geometric models of the mould and applying on its surface a pressure of 5\u00a0bar, corresponding to the internal pressure applied in the main chamber of the mould during the process. In a first case the lattice structure was computed by means of an equivalent material featuring the same macroscopic properties of the lattice, as defined by the CFD analysis (see Fig.\u00a05a). The adoption of an equivalent material to simulate the presence of the 3D lattice structure into the conveying chamber allowed reducing significantly the complexity of the model and the computational time. A second situation was then considered, where the 3D lattice was totally omitted to derive information about the contribution of the lattice to the mould stiffness (see Fig.\u00a05b). The information drawn by the FEM and the CFD models allowed a more robust design of the moulds while keeping the lightweight property as the main target. Figure\u00a06 shows an interrupted print of the designed mould made in 316L stainless steel. The complex lattice occupying the inner volume of the conveying chamber is clearly visible on the exposed section. It is to recall that the building platform of 280 \u00d7 280\u00a0mm was almost fully occupied by the printed part. Despite the optimized orientation of the part, extensive use of supports was needed to sustain the correct generation of the surfaces to be printed in the upcoming layers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002274_j.promfg.2020.02.008-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002274_j.promfg.2020.02.008-Figure2-1.png", + "caption": "Fig. 2. CAD model of the prototype of cooling system.", + "texts": [ + " From the viewpoint of safety, coolant system should be sealed because vapor of coolant may cause troubles in the laser system, and the sudden cooling may lead to metal splinters. Furthermore, coolant should be collected because the DED machine is not generally designed for the usage of liquids inside of the deposition area. Considering that the laser nozzle is difficult to cover into the seal system, the cooling system should be opened during deposition and closed during cooling. Taking the above conditions, the design of cooling system is decided as shown in Fig. 2. In addition, Fig. 3 shows the appearance of prototype system mounted in the DED machine, which was used in this study. The cooling system is shown while the sealing system is closed, which represents the \u201creadyto-cool state\u201d and the cooling system is shown in the \u201creadyto-deposit state,\u201d which means that the sealing system is open, and the laser head is in the starting position. The cooling system shown in Figs. 2 and 3 can be divided into four main elements. The aluminum framework, the pump and nozzle system, the tank and sealing cage and the cable system for opening and closing the sealing cage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000477_978-3-030-20131-9_72-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000477_978-3-030-20131-9_72-Figure3-1.png", + "caption": "Fig. 3. The PRoHep-LCT robotic system", + "texts": [ + " The number of the mobile elements is 3N (the output elements of the two modules as well as the mobile platform). The mechanism has two class 4 joints (the two cardan joints) and two of class 3 joints (the two modules). The number of degrees of freedom is calculated as follows [11]: 1 4 3 5 6 5 3 2 5i i M F N i F C M N C C (1) The second module, for the positioning of the ultrasound probe, shown in Figure 2, has a similar configuration as the first one, the two modules being assembled in a mirrored configuration on a fixed frame to enable the manipulation of both instruments in the same time (figure 3). Each of the two modules of the PRoHep-LCT robotic system has five degrees of freedom with 5 independent parameters for the mobile platform: , , , ,E E EX Y Z . For the inverse kinematic model the coordinates of the characteristic point E, , , , ,E E EX Y Z , are considered known along with all the geometric parameters of the structure, while the values of the active coordinates must be determined: 1 2 3 4 5, , , , .q q q q q Using the coordinates of the point E, the coordinates of the Cardan joints can be calculated: 1 1 1 cos( ) sin( ) sin( ) sin( ) cos( ) A E A E A E X X d Y Y d Z Z d 2 1 2 1 2 1 cos( ) sin( ) sin( ) sin( ) cos( ) A A p A A p A A p X X l Y Y l Z Z l (2) Using eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003055_s12206-020-0623-4-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003055_s12206-020-0623-4-Figure3-1.png", + "caption": "Fig. 3. Machining model for left-hand hypoid gear.", + "texts": [ + " The symbols BHR\u0394 , 0\u0394\u03b1 , s\u0394\u03b3 , \u0394\u03b5 and a\u0394\u03b3 are used to represent the spheric radius error, the pressure angle error, the rake angle error, the regrind angle error and the cutting side relief angle error, respectively. A mathematical model can be established by simulating the actual machining process with the equations of the cutting edge and coordinate transformation matrixes. The machining coordinate systems with specific details of left-hand hypoid gear for the transverse section, the cutter-plate and the processing machine are shown in Fig. 3. The expression of the major cutting edge ( pM o\u2212 ) in qcS coordinate system can be represented as follows: ( ) ( ) ( ) ( ) ( ) 0 0 2 sin / 2 sin 2 0 . 2 / 2 cos 2 1 \u23a1\u2212 \u2212 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5= \u23a2 \u23a5\u2212 \u23a2 \u23a5 \u23a2 \u23a5\u23a3 \u23a6 qr BHc BHc BHc BHc BHc BHc R u R R u R u R R \u03b1 \u03b1 (1) The expression of the fillet (M F\u2212 ) in 1S coordinate sys- tem can be represented as follows: ( )1 cos 0 sin 1 bhc bhc r r \u23a1 \u23a4 \u23a2 \u23a5 \u23a2 \u23a5= \u23a2 \u23a5 \u23a2 \u23a5 \u23a3 \u23a6 \u03b8 \u03b8 \u03b8 r (2) where the parameters u and \u03b8 are the variables for the major cutting edge and the arc segment in the transverse section, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003696_0954409720962245-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003696_0954409720962245-Figure3-1.png", + "caption": "Figure 3. Wheel/rail space contact model.", + "texts": [ + "19 Then the eight wheel profiles needed for subsequent calculations were obtained (Figure 2) The computational formula is given by y\u00f0x\u00de \u00bc f\u00f0x\u00de \u00fe e\u00f0x\u00de; x 2 \u00bda; b (1) where f (x) is cubic spline smooth curve, e(x) is the random error, and y (x) is the experimental data, [a,b] is the abscissa range of the wheel profile. Wheel-rail contact geometry is a basic problem in dynamics of a vehicle-track system. The degree of wheel/rail profile match has a direct influence on the operational quality of trains. The geometry of wheelrail contact is normally determined by deriving the x, y, z-coordinates of a random point along the trace in an absolute coordinate system for wheelset using the trace method.20\u201323 Point c shown in Figure 3 represents a wheel-rail contact point, and point c0 is the lowest point of a wheel\u2019s actual rolling circle. xO2 \u00bc dwlx yO2 \u00bc dwly zO2 \u00bc dwlz 8>< >: (2) lx \u00bc cos\u00f0hw\u00desin\u00f0uw\u00de ly \u00bc cos\u00f0hw\u00decos\u00f0uw\u00de lz \u00bc sin\u00f0hw\u00de 8>< >: (3) x \u00bc xO2 \u00fe lxRw tandw \u00fe xw y \u00bc yO2 Rw 1 l2x l2xly tandw \u00fe lzm \u00fe yw z \u00bc zO2 Rw 1 l2x l2xlz tandw \u00fe lym \u00fe zw 8>>>< >>>: (4) m \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1 l2x 1\u00fe tan2\u00f0dw\u00de q (5) Where xO2 , yO2 , and zO2 are the coordinates of the rolling circle\u2019s center, O2, in the translational coordinate system for the wheelset; Rw is the radius of the rolling circle; lx, ly, and lz are the direction cosines of the wheelset center line in the translational coordinate system; dw is the inclination angle of tread; xw, yw, and zw are the absolute coordinates of the wheelset\u2019s center of mass; hw and uw are the wheelset\u2019s roll angle and yaw angle, respectively; and dw is the horizontal coordinate of each rolling circle of wheel tread in the absolute coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001903_j.jmmm.2019.166372-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001903_j.jmmm.2019.166372-Figure15-1.png", + "caption": "Fig. 15. The schematic picture of the electrical machine under study (1/4 model).", + "texts": [ + " 14 shows the BH curves and iron losses at 50 Hz in the damaged region, which are used for the calculation of motor characteristics. This figure demonstrates significant reduction of permeability and increase of iron loss in the damaged region. The small anisotropy in Fig. 14 agrees well with the discussions in the previous chapter. The magnetic characteristics of the damaged region calculated in 4.1 is used for the simulation of an electrical machine. The schematic picture of the machine under study is shown in Fig. 15. Table 1 shows the specifications of this machine. The stator core of this machine is composed of 12 pieces. Each piece has three interlocks; two in the yoke and one in the tooth. The damaged regions with a size of 6 mm*6 mm caused by interlocking is also shown in Fig. 15. To simplify the calculation, the average data of magnetic characteristics for parallel and perpendicular directions shown in Fig. 14 were used for the isotropic characteristics in the damaged regions in Fig. 15. Fig. 16 shows calculation results of no-load characteristics for this machine. The rotation speed is 1500 rpm. Fig. 16(a) demonstrates that interlocking increases the amplitude of cogging torque in this machine. This behavior is explained by the non-linearity of the BH curve in the stator core [3]. Fig. 16(b) demonstrates that interlocking increases iron loss under no-load condition. Both results clearly show that it is important to consider the influence of interlocking for accurate calculations of motor characteristics" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure7-1.png", + "caption": "Figure 7 First-order mode", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0001707_icems.2019.8921756-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001707_icems.2019.8921756-Figure1-1.png", + "caption": "Fig. 1. The model of the tooth and stator core.", + "texts": [ + " Finally, a damped sandwich structure is used to weaken electromagnetic vibration. II. MODAL ANALYSIS The calculation of the modal analysis on the stator core is of great importance. Here, we only consider the radical vibration of the models. This part, the IPMSM motor which has 6 pole-pairs is considered. The slots shape of the stator core are pear-shaped, which are a kind of semi-closed slots. The parameters of the stator core are shown in Table I. The FEM model of the stator core and model of stator tooth are shown in Fig. 1. The material of the stator is stainless steel. The density of the stator is 7600kg/m3, the elastic modulus is 120GPA,the poisson's ratio is 0.27.The element size of the model is 10 mm. The stator core modal can be simulated by Workbench, the Tab. II shows the natural frequencies of the second mode to the fourth mode. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE The modal analysis of the stator had closed relevance to the structure parameters. Different sizes of the stator core can affect the trend of the natural frequencies, therefore, we can study the stator structure to seek for what factors which may have influences on the natural frequencies" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002834_tia.2020.2998667-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002834_tia.2020.2998667-Figure1-1.png", + "caption": "Fig. 1. Vector expression of the machine power interruption. (a) Normal operation. (b) Stator current decreasing after power is interrupted. (c) Fully power-off", + "texts": [ + " Stator currents will decrease in few milliseconds through the diodes inside the IGBT modules. The rotor currents, however, will decrease at a much slower pace because the rotor is shortcircuited, and only the rotor resistance dissipates the energy. Traction systems like locomotive trains have huge inertia. Therefore the rotor is continuously running with little speed loss after cutting off the PWM pulse. According to the law of electric-magnetic induction, the rotating air-gap flux induces a back-EMF on the stator. Fig. 1 presents the above process in the vector expressions. Fig. 1(a) describes the rotor and stator flux linkage while the power supply is on. I\u0307s is the stator current, and I\u0307r is the rotor current vector. I\u0307g is an equivalent current vector to excite the air-gap flux linkage vector \u03c8\u0307g . The air-gap flux vector \u03c8\u0307g plus the rotor leakage flux \u03c8\u0307lr is the rotor flux \u03c8\u0307r. Similarly, the stator flux vector \u03c8\u0307s is synthesized by \u03c8\u0307g and the stator leakage flux \u03c8\u0307ls. Whenever the power supply is cut off, the stator current sharply drops to zero. The rotor flux should remain unchanged according to the Lenz Law. Therefore, additional rotor current is induced. In the period of I\u0307s dropping, as shown in Fig. 1(b), the amplitude of I\u0307s decreases, and its phase keeps constant. In the rotor side, the q-axis component of the rotor current vector starts to decrease, and the d-axis component increase from zero. Eventually, in Fig. 1(c), the stator currents have been reduced to zero completely, and the rotor current vector is aligned with the rotor flux vector. The air-gap flux vector is also aligned with the rotor flux vector as well. Meanwhile, the amplitude of the rotor flux and the rotor current decreases with the rotor time constant. Voltage and flux mathematical equations of induction ma- Authorized licensed use limited to: University of Exeter. Downloaded on July 16,2020 at 01:52:59 UTC from IEEE Xplore. Restrictions apply" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000609_978-3-030-12684-1_17-Figure17.14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000609_978-3-030-12684-1_17-Figure17.14-1.png", + "caption": "Fig. 17.14 The geometry of the lattice parts within the outer walls", + "texts": [ + " The printing of lattices, which were aligned with the build direction to avoid using support material, introduces very different anisotropies. It is posited that the anisotropy here is dominated by lattice geometry, not build layer orientation (as is the case for solid parts). The stiffest lattice orientation is upright, since axial loads are resisted with continuous material. The DB orientation is moderately less stiff than upright, and the thin orientation is the least stiff. Excluding the external walls of the parts (a common element between all three), the DB part has better axial connections than the thin part (Fig. 17.14). If we were to test the in-plane bending modes, we expect the thin and upright parts to be stronger than the DB part for the same reason. The main result of this work with respect to torsion modes and modal damping ratios is that the DB orientation has the highest torsion frequency, as well as the most damping in both torsion and out of plane modes. If one imagines an oscillating AM part as laminated sheets rubbing against each other, the DB part has by far the largest sheets, and so the highest friction loss" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002376_b978-0-08-102941-1.00017-1-Figure17.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002376_b978-0-08-102941-1.00017-1-Figure17.1-1.png", + "caption": "Figure 17.1 A conventional approach for modeling lattice structures. Estimates of bulk material properties for a specific fabrication process serve as input to unit cell models; effective or homogenized unit cell properties serve as input to large-scale lattice structure models. No process-induced variability is explicitly modeled.", + "texts": [ + " Implementation of the strategy begins by investigating the repeatability and predictability of the mechanical properties of periodic octet truss lattice structures fabricated with DMLS and loaded in compression. Then, variability in strut-level mechanical and geometric properties is characterized experimentally, and the results are utilized to enhance the predictability of the mechanical properties of periodically repeating lattice structures. 17.2 A strategy for predicting the mechanical properties of additively manufactured metallic lattice structures via strut-level mechanical property characterization The conventional approach to modeling lattice structures is illustrated in Fig. 17.1. In this approach, the design engineer assumes that the AM process fabricates bulk material that is arranged into unit cells, which are organized into lattice structures. The mechanical properties of the bulk material are characterized using standard ASTM testing procedures to evaluate Young\u2019s modulus, yield strength, ultimate strength, elongation at break, and other mechanical properties for the base material (e.g., 17-4PH stainless steel, Ti-6-4) and process of interest (e.g., DMLS, SLM). The material properties could be generically agnostic to processing parameters, or they could be distinguished by build orientation, location on the build chamber, or other processing parameters of interest to the designer", + "2(d)), which is adopted in this research study, relies on building experimental models of the effective mechanical properties of struts and utilizing those empirical models as direct input to computational models of an overall lattice structure. This approach is more computationally efficient than other approaches because it does not require any simulation of fine-scale features. Significant experimentation is required, however, to build the empirical models of strut behavior. This approach is implemented in Sections 17.5 and 17.6, where it is found to estimate the mechanical stiffness of lattice structures with an order of magnitude less error than the standard approach depicted in Fig. 17.1. In summary, the focus of this research is to demonstrate the approach depicted in Fig. 17.2(d) and contrast it with the conventional approach illustrated in Fig. 17.1. The conventional approach is implemented in Sections 17.3 and 17.4. Specifically, Section 17.3 describes fabrication and testing of lattice structures and standard tensile bars. Section 17.4 reports the results of FEA of the lattice structures based on the bulk properties derived from the standard tensile bars and compares those lattice structure properties to the experimental results. The approach depicted in Fig. 17.2(d) is implemented in Sections 17.5 and 17.6, which describe experimental testing of strut-level tensile specimens and incorporation of the experimentally derived effective mechanical properties of the struts into a more refined FEA of the lattice structures" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003121_0954405420932445-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003121_0954405420932445-Figure2-1.png", + "caption": "Figure 2. Geometry construction of miniature bevel LG.", + "texts": [ + " Where u1, n and t are consistent with the parameters in equation (1). Equations (3) and (4) correspond to equation (1), which describes the driving contact curve R1, while the driven curve R2 has the same form x2 + y2 sin2(u1) = z2 cos2(u1) \u00f03\u00de x= n sin (u1)3 t3 cos (t) y= n sin (u1)3 t3 sin (t) \u00f04\u00de In step 1, the enveloping conic surface of the miniature bevel LG described in equation (3) needs to be manufactured using the layer-by-layer method, as shown in Figure 1. In step 2, the material other than the teeth on the conic surface continues to be removed. Figure 2 shows the geometry constructions of the driving and driven miniature bevel LGs assembled. The dashed area indicated by diagonal lines is the tooth space from where the material is removed. In brief, a micro circular cone is manufactured in the first step of the process; gear teeth are manufactured in the second step of the process. This is named the twostep method of NSPLA. To get an intuitive understanding of the proposed two-step method, please refer to the supplemental material. The two-step method of NSPLA has two advantages: First, the scanned area conforms to the geometric features of the structure, which simplifies the scanning process; second, it is easy to characterize the shape errors of the conical helix, which is decomposed into the shape error of the enveloping cone and Archimedes screw" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000325_978-981-13-5953-8_48-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000325_978-981-13-5953-8_48-Figure1-1.png", + "caption": "Fig. 1 Structure of a two-wheeled mobile robot", + "texts": [ + " The problem space for the multi-robot navigation is considered as a two-dimensional world map. The environment consists of static obstacles of different shapes and twowheeled multiple mobile robots. Every robot in the environment has predefined start and the goal position. The objective of the problem is to find an optimal and safe path for the individual robot from their respective start position to the goal position. In the initial setup, every robot\u2019s orientation is toward the goal position. The structure of a two-wheeled mobile robot is shown in Fig. 1. The current position of a two-wheeledmobile robot in two-dimensional coordinate system is denoted by (x, y) and orientation of the robot with respect to abscissa is \u03b8 . All the robots in the environment are of identical structure with radius r . The angular and translational motion of the robot is denoted by \u03c9 and v, respectively. In order to navigate safely, every mobile robot generates successive waypoints by avoiding obstacles in their way. Let {p1, p2, . . . , pm} be the \u201cm\u201d waypoints generated by the mobile robot while navigating from source to goal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001985_robio49542.2019.8961540-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001985_robio49542.2019.8961540-Figure3-1.png", + "caption": "Fig. 3 The joint layout of the first segment for the equivalence manipulator", + "texts": [ + " Kinematic equivalence means that the kinematics model (position and direction vector of the end of any segment) of a new hyper-redundant manipulator which is be constructed by simplifying and rearranging joints in each segment is equivalent to that of the developed hyper-redundant manipulator shown in Fig. 1. Similar to the original hyper-redundant manipulator, each segment of the kinematic equivalence hyper-redundant manipulator is also consist of the same joint layout. Hence, taking the first segment of kinematic equivalence hyperredundant manipulator for instance, the specific joint layout only is (Roll-Yaw)-Yaw-Yaw\u2026Yaw-Yaw that is shown in Fig. 3, whether n is even number or odd number. Furthermore, there are also n joints for each segment of the kinematic equivalence hyper-redundant manipulator. And the first is universal joint which is active joint and is formed by Roll axis and Yaw axis. The others are revolving joints that are passive joint rotating in the direction of Yaw axis. The rotation angle of each passive joint is equal to that of the Yaw axis of the first joint, i.e., 1,1 1,2 1,3 1,n 1 1,n (4) As shown in Fig. 3, there are n+1 rotation axes for each segment of the kinematic equivalence hyper-redundant manipulator. For the ith segment, the movement angle of the first rotation axis can be denoted by i. And the symbol i is the rotation angle of the second rotation axis. In addition, the movement angles of the others which are moving with the second rotation axis synchronously can be also written as i. Hence, the homogeneous transformation matrix from the end of ith segment to the end of (i-1)th segment can be expressed as 1 1 1 1 , , , 0 0 0 1 0 0 0 0 1 i i i i i i i i i i i i i i i i i i i i i n n n j i j n n n j i i i j n n n j j i j n Rot z Rot x Trans z L c s c s s s s L s c c c s c s L s c c L T R p (5) According to (5), the pose of the end-effector of the kinematic equivalence hyper-redundant manipulator is yielded by recursion law, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001906_s12206-019-1203-3-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001906_s12206-019-1203-3-Figure2-1.png", + "caption": "Fig. 2. Strain gauge locations on beam four of the proposed elastic structure.", + "texts": [ + " Forces acting on the sensor in the x or y directions result in larger deformations of the vertical beams (V-beams), whereas forces acting in z the direction result in larger deformations of the horizontal beams (H-beams). To obtain high strain values and reduce the overall height of the elastic structure, the V-beams contain halfcircle cutouts that increase the stress concentrations on the surfaces of the V-beams at the tops of the cutouts. All candidate sensing points are specified for each T-shaped beam, as shown in Fig. 2, where the subscript i in Sij indicates the beam number (1\u20134) and the subscript j indicates the strain gauge location (1\u20138). For example, S41 is the sensing point on beam four at location one. The second subscripts two, four, and six in Sij specify bottom surface points. There are four sets of candidate strain detection points for the four T-shaped components. In this design, the H-beams of the T-shaped components are machined in the lower part of the elastic structure and the Vbeams are joined to the upper part", + " 3(a)-(f) correspond to the finite element method (FEM) results for the H-beam when a single force in the x, y, and z directions, and single moment in the x, y, and z directions were applied to the sensor. Figs. 4(a)(f) correspond to the FEM results for the V-beam when a single force and moment in the x, y, and z directions were applied to the sensor. For this analysis, we used 200 N for the single applied force and 10 Nm for the single applied moment because the target maximum force and moment for sensor development were specified to be \u00b1200 N and \u00b110 Nm. For each T-shaped beam (including V-beam and H-beam), 8 candidate points (refer to Fig. 2) are selected in advance for strain gauge bonding and strain measurement. Then for given maximum force or maximum torque in each direction, surface strains for all 32 candiate points (8\u00b44) are computed according to the thicknesses of the V-beam and H-beam by DOE analysis, and the most efficient point pairs with big strain differences are determined. Figs. 3 and 4 are typical six strain plots with relatively big strains by DOE analysis. As shown in Figs. 3 and 4, the surface strains at various detection points vary significantly according to the thickness of the H-beam, but vary only slightly according to the thickness of the V-beam" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003917_tmech.2020.3036571-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003917_tmech.2020.3036571-Figure1-1.png", + "caption": "Fig. 1. Experimental testbench, assembled at NASA JSC, comprising (a) actuator base link (brushless dc motor, harmonic drive transmission, and rotary spring obscured), (b) actuator output link which rotates relative to the base link, (c) spring deflection sensor, (d) calibrated torque sensor, and (e) adjustable output inertia (these two weights are tightly lashed to the output lever using high-strength Vectran cable).", + "texts": [ + " Though we do not explicitly treat it here, this problem shows up indirectly even in the multijoint case [18], [19]. We contribute the following: 1) necessary and sufficient gain conditions for nominal system passivity; 2) gain upper bounds derived from stability and minimum phase aspect of passivity with time delay and derivative filtering; 3) a phase-relaxation of passivity that is related to the regenerative efficiency of springs; 4) sufficient conditions for nominal relaxed passivity; 5) a hardware example (see Fig. 1) of a stiffer-than-passive behavior. Our strategy uses two models, as shown in Fig. 2. The nominal model is a simplified model that allows for analytical pole and zero placement in the design process. The design model is more complete but less tractable and is used to justify bandwidth limits on the pole/zero placement process from the nominal model. In both models, there are two nondisturbance inputs (motor torque \u03c4m and joint torque \u03c4j) and two (linearly independent) outputs (\u03b8m and \u03b8s). These outputs can construct a third output \u03b8j which is part of the joint port", + " phase condition guarantees that this feedback results in a nonnegative phase margin. Note that these conditions only ensure \u03c8-relaxed-phase passivity at the nominal model\u2019s joint port. An analytical expression for \u03c8-relaxedphase passivity of the design model\u2019s joint port is beyond the scope of this article. However, if (11) and (12) hold, the relaxed phase condition can be graphically verified after the fact using the Bode plot of the design model\u2019s joint port\u2019s integral admittance (see Fig. 3). Using a NASA Valkyrie actuator (Fig. 1, [26]), we demonstrate the behavior of the controller from Fig. 3. We identify the parameters of our linear model (Table I) using closedloop tests driven by the joint output (and not by the motor, as is more typical). This testing scheme makes use of the assumed model structure for an SEA. The first test has no feedback controller. From the empirical integral admittance of the motor, \u03b8\u0302m/\u03c4\u0302s, we find a linear estimate for the reflected motor dampingBm. The second test finds the negative motor velocity feedback which corresponds to the boundary of stability using bisection search" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure6.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure6.10-1.png", + "caption": "Fig. 6.10 Advanced technology features of LCA Aircraft", + "texts": [ + " It was the only helicopter (from among Indian, European,American andRussian origin helicopters in use in India formilitary and civil transport) which could fly at the high altitude under adverse weather (rain and storm) conditions and rescue disaster victims during the \u2018Uttaranchal\u2019 landslide in June/July 2013. This led to Indian team receiving the prestigious world award for the rescue and relief operation from the American Helicopter Society. Light Combat Aircraft (LCA) is another major achievement for the aerospace and materials Scientists/engineers. Figure 6.10 illustrates some of the advanced technology features of LCA. It is designed as a highly agile, and worlds lightest advanced technology multi-role combat fighter in the empty weight category of 6000\u20137000 kg and a speed of 1.5 M and a service cealing of 16 km. In both these above projects weight optimization was aimed during the design and prototype development/manufacturing phases to achieveminimumweight maintaining structural integrity and the high performance requirements. Principles and processes as outlined earlier to design with alloys of high strength to weight ratio, provision of lightening holes/cut outs, integral milling, local thinning, etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003172_042002-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003172_042002-Figure1-1.png", + "caption": "Figure 1. Underwater Vehicle Movement Design", + "texts": [ + " Some other application examples of using this vehicle can be found in [6], [12]\u2013[14] . This paper proposes the mathematical model of ROV movement. Combination of statics disciplines and dynamics of ROV movement are needed to formulate the mathematical model of ROV. Statics refers to the forces and moments that work on the physical system around the equilibrium point, dynamics refers to the effect of force on the movement of the object. The most important thing in designing this underwater vehicle is how to formulate the movement of the vehicle. Figure 1 shows a picture of the design and position of the vehicle on the coordinate plane. Based on these coordinates, there are 6 (six) degrees of freedom (DOF), where the vehicle can move as follows: forward, surge sideways (sway) and float up (up) in the direction of the axis X, Y and Z and roll, go up and turn (yaw) following the X, Y and Z axis. The vehicle weight is assumed that capable of making the entire vehicle sink, so control will be prioritized to handle the downward force (buoyancy). The center of gravity is in the middle of the vehicle's body, so this will create a robotic stability during pitch and roll movements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001569_ecce.2019.8912739-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001569_ecce.2019.8912739-Figure5-1.png", + "caption": "Fig. 5 Centroid calculation", + "texts": [ + " Each pole of the rotor is considered as embedded with the adjacent poles at the air gap iron bridges (K1 and K2), and with the rotor body at the central iron bridge (K3) (Fig.4). The centrifugal force Fcentrifugal is the main source of stress within iron bridges. It is applied at the point G, the center of gravity of the active part of mass m of the rotor which is defined as the upper part of the rotor pole including all the magnets. G is calculated by dividing the active part into different geometric shapes, then taking the sum of the mass of the shapes, multiplied by their positions, and divided by the total mass of the active part (Fig.5). 2\u03a9\u00d7\u00d7= Glcentrifuga RmF (1) Where RG is the radius of the centroid point G and \u2126 is the rotational speed of the rotor. The FEA of the rotor deformation due to the centrifugal force, shows that: beams 1 and 2 undergo a flexion stress and beam 3 a traction stress. This allows to obtain the effort configuration shown in Fig.6. It has been shown, based on the same previous FEA, that the relationship between the centrifugal force and these efforts is linear and is described by constant coefficients (A, B and C)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-Figure6-1.png", + "caption": "Fig. 6. Overview of linear mechanism", + "texts": [ + " When the pin 2 leaves the L-shaped groove, the node N2 can perform linear motion freely with respect to the node N3. \u2022 STEP4; The node N2 can perform rotation and linear motion freely with respect to the node N3. We can extend it sequentially by repeating this process. Contraction of the nodes is possible by performing in the order of STEP4 to STEP1. Note that there is a mechanism that releases pin 1 by releasing a certain amount of force in the extension direction. Please refer to the appendix for details of the mechanism. 3) Mechanical implementation: Fig. 6 shows an overview of the designed linear mechanism. The linear mechanism is driven by three motors. The motor M1 performs a linear motion of the square shaft S1, S2. The motor M2 performs a rotation of the square shaft S1 with respect to square shaft S2. The motor M3 performs a rotation of the square shaft S1, S2 with respect to the base. This rotation is used to release the constraint of linear motion by the constraint mechanism. The motor M1 drives the slide screw and the square shaft S1, S2 move linearly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001144_s1068798x19090168-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001144_s1068798x19090168-Figure4-1.png", + "caption": "Fig. 4. Transceiver housing for active phased array antenna produced by selective laser melting (a); and tomographic data for the housing produced (b). In tomography, the dark areas correspond to empty space.", + "texts": [ + " The test results indicate that the physicomechanical properties of the material are acceptable for the production of the required parts and hardly differ from those obtained by traditional technology. The porosity of the material is no more than 1%. The thermal conductivity in the printing plane is 175 W/(m K); in the direction of growth, it is 130 W/(m K). The specific heat is 890 J/(kg K). The Young\u2019s modulus is 80 GPa, that is higher than the standard value for aluminum (70 GPa). The yield point is 160 MPa in extension and 270 MPa in compression; the tensile strength is 255 MPa; and the surface roughness after shot treatment is Ra = 5.5 \u03bcm. In Fig. 4a, we show the housing produced. A tomographic image is presented in Fig. 4b. The housing geometry is satisfactory, while the cooling channel is practically free of powder. Tomographic data reveal only one point with powder adhesion at the wall (the vertical section of the channel ahead of the output). In ENGINEERING RESEARCH Vol. 39 No. 9 2019 terms of distortion, it is important to note the arched geometry of the lower housing face, which is close to the platform in printing (Fig. 4a). These geometric errors are associated with significant residual stress in the sintered material during the manufacturing process. On account of the size of the sintered cross section, relatively large distortion appears there. This distortion may be eliminated by appropriate selection of the manufacturing conditions and the orientation of the product in the working chamber. That will be considered in future research. (1) By additive technology, we may manufacture complex parts unsuitable for production by traditional methods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002935_0954408920932358-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002935_0954408920932358-Figure2-1.png", + "caption": "Figure 2. The one element radiation structure and the two element radiation structure.", + "texts": [ + " Inspired by the chemical molecular structures which have good stability and space symmetry, this paper begins to study those structures including methane and pentane. The zero element radiation is not a completely symmetric structure and it is a simple model, so many studies put it as an initial model to study the power generation of pipeline. However, it has no real value in itself. Therefore, this paper mainly studies the one element radiation and the two element radiation. Their dynamic analysis will be carried out in Figure 2. In Figure 2 above, it is one element radiation structure with four groups of pipes converging at one point and diverging in all directions, and their ends are tetrahedral, similar to the molecular structure of methane; the following figure is based on the above, connecting three groups of pipes at the end of four groups of pipes respectively, and the structure is similar to the molecular structure of neopentane. Meanwhile, Figure 2 has established the coordinate system and marked the relevant parameter information. Solve the equations of induced electromotive force The design of energy collector is based on Faraday\u2019s electromagnetic induction law. When the magnet passes through the coil, the energy will be generated in the coil. In fact, the study of all the power generation devices can be illustrated by analyzing the single one element pipeline, because the energy collecting device is symmetrical in structure and it has the same principle as the two element pipeline model", + " Pipe length Coil width Coil diameter Coil position Spring stiffness coefficient 500mm 20mm 1mm 250mm 10/20/40/80 N/m Resistor of coil (R0) Mass of magnet (m) Length (L) Friction Coefficient (l) 11X 50 g 500mm 0.3 Therefore, with the analysis of two factors, spring stiffness coefficient will be set 20N/m and the rotation speed will be limited range from 30 rpm to 120 rpm. Just studying the angle between the pipeline and the assumed rotation plane can give the comprehensive overview of the motion characteristics and the power generation efficiency. As shown in Figure 2 the rotation plane is set as the YOZ plane, and the OX is the rotation axis. The Z axis is perpendicular to the paper. Whatever the angle of inclination is, it always produces a fixed angle a with the YOZ plane when it rotates around the OX axis. When the rotation speed is set up the 60 rpm, the angle between the pipe and the rotating plane is 0/30/ 45/60 degrees respectively, and then the motion characteristics are simulated in turn. The results of speed are shown in Figure 6. For any single pipe unit, the permanent magnet is affected by spring force, tube wall friction, gravity and the centrifugal force", + " When the angle between the pipeline and the rotating positive direction are 0/30/45/60 degrees respectively, the kinematic characteristics at different rotational speeds (30/60/90/120 rpm respectively) are carried out by Adams. The results are shown in Figure 9. Figure 9 presents the velocity characters of different angles. With the increase of angle, the amplitude of the velocity decreases. Moreover, good stability is obtained. The two element structure is similar with the one element structure, but it has the binary part while the later only has the unitary part. For the sake of presentation, they are named the unitary pipeline angle and the binary pipeline angle in the two element structure as shown in Figure 2. Following will analyze the relationship between the two element pipeline and generating capacity. The unitary pipeline angle is 0/30/45/60 degrees respectively, and then the four small pictures divide the binary pipeline angle into 30/45/60/90 degrees in turn as shown in Figure 10. From Figure 10, when the unitary pipeline angle is 0/30/45/60 degrees respectively, the maximum EMF generated is about 3.2/2.7/2.3/1.8 V. With the increase of speed, the amplitude of EMF decreases gradually. It\u2019s smaller than the one element structure, but it\u2019s more stable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003766_j.triboint.2020.106684-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003766_j.triboint.2020.106684-Figure7-1.png", + "caption": "Fig. 7. The bearing lubricating characteristics for different bearing speed cases. (a) The oil volume fraction of bearing flow field for rated speed, (b) oil volume fraction of bearing cross section for rated speed, (c) velocity field of bearing cross section for rated speed, (d) oil volume fraction of bearing flow field for half rated speed, (e) oil volume fraction of bearing cross section for half rated speed, (f) velocity field of bearing cross section for half rated speed, (g) oil volume fraction of bearing flow field for state with zero speed, (h) oil volume fraction of bearing cross section for state with zero speed, and (i) velocity field of bearing cross section for state with zero speed.", + "texts": [ + " More understanding of the key parts of lubricating characteristics of lubricating system of WPG is discussed in this section. In WPG, since the output shaft front bearing has a faster speed, it needs a better lubrication condition. Thus, the lubricating characteristics of output shaft front bearing is studied. The rated speed of bearing is 1760 r/min. According to the flow distribution analysis results of WPG lubricating system, the oil flow rate for lubricating the bearing is 17.12 L/min under the lubricating oil input flow rate case of 140 L/min. Fig. 7 depicts the bearing lubricating characteristics for different bearing speed cases. Fig. 7(a)\u2013(c) give the lubricating characteristics for rated speed case. Fig. 7(d)\u2013(f) show the lubricating characteristics for half rated speed. Fig. 7(g)\u2013(i) plot the lubricating characteristics for the state with zero speed. As in Fig. 7(a) and (d), the air-oil distribution inside the bearing flow field is uneven one, the oil volume fraction of position at the front of nozzle is higher than those of other positions. The results show that the bearing speed has a great influence on the oil volume fraction. Fig. 7 (b) and (e) depict that the air-oil distribution in the radial direction of bearing is also uneven one, which is caused by the centrifugal force. Since the denser oil phase will fly to the outer raceway, the oil volume fraction of outer raceway is higher than that of inner raceway. Fig. 7(c) and (f) show that the lubricating oil splashes outward due to the centrifugal force. The performance of oil splashes outward is more obvious with the increment of speed. Fig. 7(g)\u2013(i) give that the lubrication effect of bearing is ineffective under the condition with zero speed. The lubricating oil is only distributed near the nozzle location. A one-dimensional fluid analysis model is presented to analyze the flow distribution of lubricating system. The effect of lubricating oil input flow rate on the flow distribution of each part in the lubricating system are studied. To analyze the bearing lubricating characteristics in the WPG, a fluid-structure coupled simulation model is presented, which is based on the CFD method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003193_012062-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003193_012062-Figure2-1.png", + "caption": "Figure 2. Pairing slopes with a semicircular channel bottom: M\u043e is channel slope laying coefficient; D\u0440 is rotor diameter; \u041d is channel depth", + "texts": [ + " On the channel with reinforced slopes, after cleaning their machines, cross-sectional profiles can be obtained that get a semicircular outline having the smallest friction surface (Figure 1, b) [9, 10]. At the same time, lateral slopes are often used to smoothly interface the lower part of the channel with the semicircular channel, they are pivotally connected to the body of the working body, which allows you to adjust them with different revelations and obtain a profile of the cleaned channels of the correct shape (Figure 2) [11]. CONMECHYDRO \u2013 2020 IOP Conf. Series: Materials Science and Engineering 883 (2020) 012062 IOP Publishing doi:10.1088/1757-899X/883/1/012062 During the overhaul of canals with non-reinforced slopes after cleaning them with a multi-bucket working body of cross digging in two passes, a stable section is also formed. It has been established that siltation of channels is mainly caused by low speeds of the water moving in them and it occurs due to the destruction of slopes and accumulation of soil deposits at the bottom of the channel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure7.13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure7.13-1.png", + "caption": "Figure 7.13 Diagram of deflections and bending moments.", + "texts": [], + "surrounding_texts": [ + "The substitution of (7.6.14) into formula (7.6.20) leads to the divergent series, similar to (7.6.15), the regularization of which can be carried out according to the formula (7.6.18). For example, for the square plate the bending moment is described by the expression M(x, y) = 16 a2 \u221e \u2032\u2211 m=0 \u221e \u2032\u2211 n=0 {(\ud835\udefc2 m + \ud835\udf08\ud835\udefc2 n)q(m,n) + (\ud835\udf08 \u2212 4)(\ud835\udf08 \u2212 1)(\ud835\udefc2 m + \ud835\udefc2 n) \u00d7 [(\u22121)m\ud835\udefc2 nA(n) + (\u22121)n\ud835\udefc2 mA(m)] \u2212 signk0 \u2212 (\ud835\udf08 \u2212 4)[(\u22121)mA(n) + (\u22121)nA(m)]} cos\ud835\udefcmx cos\ud835\udefcny (\ud835\udefc2 m + \ud835\udefc2 n)2 + signk0 . (7.6.21) As an example let us examine a square plate with the given size of 2\u00d7 2, with the following types of load: 1. Force concentrated in the center of the plate P = 1; 2. Uniform load q1 = 100, distributed along the area of the plate with size 0.1\u00d7 0.1 in the center of the plate. 7.6 The Forced Oscillations of a Rectangular Plate 427 The diagrams of the static deflections and bending moments in the sections throughout the diagonal x = y and along the axis x = 0 are obtained by taking into account the 25 terms of the series, and are shown in Figures 7.13 and 7.14. Moreover, with the last version of the load of the value of deflection w = 0.272 and bending moment M = 0.2987 in the center of plate differ from those obtained by an alternative method in the work of V.I. Travush, V.K Sangadzhiev [364] by values less than 1%. The analysis of the dependence of the deflections of the plate on the frequency of load was conducted for the last version of load. An increase in the frequency of load leads to the decrease of the given rigidity k0(\ud835\udf14), and consequently to an increase in the absolute value of reduced length l. At the same time the sign k0(\ud835\udf14) changes in the transition into the \u201ctransresonant\u201d regime (\ud835\udf14> k(\ud835\udf14)/m)." + ] + }, + { + "image_filename": "designv11_80_0001074_ccdc.2019.8832404-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001074_ccdc.2019.8832404-Figure6-1.png", + "caption": "Fig 6. Meshing Model of Motor Rotor", + "texts": [ + " If the mesh is properly split, the accuracy of computer processing can be improved, the simulation results can be reduced, and the actual errors can be made, so that the results of the simulation calculation are more in line with the true values.[9] Because the software automatically splits the mesh sparsely, it will reduce the accuracy of calculation. Therefore, when using 3D finite element software for simulation, manual meshing is required. The meshing model of the double stator permanent magnet spherical motor rotor is shown in Fig. 6. Define the direction of coercive force of the permanent magnet of the motor and simulate the magnetic field. The motor adopts a structure of a pair of coils inside and outside the rotor. Therefore, the magnetic field generated by the permanent magnet of the rotor of the motor is the main component of the magnetic field, and the influence of the air-core coil on the magnetic field of the permanent magnet is negligible. Therefore, when the magnetic field is analyzed by the 3D finite element method, the magnetic field inside and outside the magnetic pole generated by the permanent magnet of the rotor is mainly modeling and analysis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure90.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure90.10-1.png", + "caption": "Fig. 90.10 Von Mises stress distribution in the debond model", + "texts": [ + " At debond regions, the internal pressure load applied on liner will not be effectively transferred to composite overwrap. This occurs because of improper bonding and insufficient bond pressure between liner-composite overwraps. Hence 1082 R. Pramod et al. a void created in-between will allow liner to consume pressure load more than its capability, which will allow it to deform excessively as it cannot effectively transfer the load to composite overwrap. Peak Von Mises stress of 293 MPa (refer Fig. 90.10) is seen at the debond locations. The obtained stress of 293 MPa is 3 MPa more than the ultimate tensile strength of AA 6061-T6 proving that liner will fail in presence of debond. Graphical representation of Von Mises stress along the axial length of liner is shown in Fig. 90.11. Figure 90.12 shows the maximum deformation (7.93 mm) in liner at debond location. This plastic deformation causes 90 Finite Element Analysis of Potential Liner Failures \u2026 1083 1084 R. Pramod et al. additional problem, as during consecutive pressurization-depressurization cycles, it may cause liner to buckle; hence, bringing down the efficiency of liner and possible failure during operation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003427_012038-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003427_012038-Figure1-1.png", + "caption": "Figure 1. The metallographic results. The tested part is in the middle, while the two sample is on its sides. Below the macro photos of the part, is the optical micrographs showing the worst cases in the vertical and horizontal sample, and the second worst case on the horizontal sample.", + "texts": [ + " Today the 3D imaging devices like computer tomography [10] are more common and can help greatly during the investigation of the worst case in the product. Therefore, the result of a metallographic testing and a tomography is compared to show difference and parallelism of the two method through the evaluation of the microporosity of die cast aluminum part. An aluminum housing was chosen which was produced by die casting. The raw material of the casting is EN-AC47100. The nominal composition of the alloy is: 10.5-13.5w/w% Si, 0.7-1.2w/w% Cu and max. 1.3w/w% Fe, beside the aluminum. The housing and the typical microstructure are shown by Figure 1. The pore structure is evaluated on the micrographs (Figure 2). Based on the tomographic image, the worst place and position of the metallographic samples are determined. The metallographic samples were cut by sawing. The examined surface of the sample was grinded mechanically by SiC particles, then polished by 3 and 1 \u00b5m diamond particles. The polished surface was etched by immersion to HF solution. Micrographs were taken by a Zeiss AxioImager M1m optical microscope. The micrographs were analyzed by computational image analysis with a self-developed Cprob software", + " Then the pores are indexed, and the distance of the outermost points is measured in the pores as the length. Additionally, the distance between the nearest pores also measured. The nearest pores are determined by a SKIZ operation [4], and the closest point of the nearest pores are measured as a distance. Some evaluations a distance limit is determined, and closest pores have to be handled as one big pore. In this case the length of the big pore is the sum of the pores and distances affected by this evaluation. The length of the pores and the images are compared. Figure 1. shows the examined housing with the position of the samples. Two samples were cut from the part for metallographic testing. The positions of the samples are selected by the tomographic image. Sections with large pores are chosen. The samples were prepared for optical microscopy, and micrographs were taken. Figure 1. also shows some micrographs where the typical microstructure also can be studied. The nominal composition is close to the eutectic composition of the Al-Si alloy system. According to it a fully eutectic microstructure can be seen on the micrograph. The eutectic structure is fine, and the silicon is not a plate like in the eutectic. The eutectic structure is modified to achieve a better mechanical property. The modification turns the silicon into small plates and globular particles. All micrograph contains micropores, which are the dark objects with different shapes", + " The globular objects like spherical shapes are probably gas inclusions. The others are micropores formed during the shrinkage in the solidification process. Figure 2. shows typical sections of the tomographic reconstructed 3D images, measured with an YXLON FF35 MicroCT. The parameters of the test are summarized in Table 1. As it can be seen, the background of the images is dark while the aluminum material is bright, so the micropores \u2013 microcracks are visible also in dark color. The image shows a top-view section, and a section according to Figure 1. 12th Hungarian Conference on Materials Science (HMSC12) IOP Conf. Series: Materials Science and Engineering 903 (2020) 012038 IOP Publishing doi:10.1088/1757-899X/903/1/012038 On both sections the micropores are visible and can be measured by the software package of the tomography equipment. The micropores on the sequence of optical micrographs are measured by the mentioned image analysis software. As it shown the pores are dark in the bright field illuminated images. The segmentation of the pores can be made by the intensity level of a grayscale image (Figure 1 and Figure 3). The segmented objects, the pores can be evaluated. The main measure of the pores is the length. The area, especially the area fraction of the pores on a section is also important as an impact to the static mechanical properties, but the fracture mechanics determines a maximal length based on the planned stress state of the part. This length basically the diameter of the plane section of the pores in case of a general single pore. But in a case of clustered pores the length is different" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000452_ilt-10-2018-0397-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000452_ilt-10-2018-0397-Figure4-1.png", + "caption": "Figure 4 The diagrammatic sketch of wheel rolling Figure 5 The changing trail of the contact patch", + "texts": [ + " If the elastic deformation occurs at this moment, P1 came into contact with P2. After interval Dt, to time t, P1 and P2 were separated and there were some differences between wheel and rail contact position along running direction at that time. Consequently, the relative displacement was emerged. The existing contact tangential force between nodes in wheel and rail contact patches included the lateral tangential force (Fcx) and the longitudinal tangential force (Fcz), and it was the main factor of wheel-rail contact fatigue and damage (Figure 4). The research emphasis of this paper was the longitudinal contact tangential force, whichwas the contact friction. The wheel-rail friction work at the wheel-rail contact patch was obtained basing on the physical meaning of work, which was the product of the node relative displacement and the contact friction. It was supposed that the state of the wheel-rail contact patch keep stable along running direction of the train. When the wheel passed through the distance of a contact patch, the changing trail of the contact patch on the wheel surface was shown in Figure 5, where the arrow pointed to the running direction of the train" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002276_edpc48408.2019.9011935-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002276_edpc48408.2019.9011935-Figure5-1.png", + "caption": "Fig. 5. Breakdown voltage test rig", + "texts": [ + " The breakdown voltage setup consists of a U-shaped plastic socket in which an L-shaped block of carbon steel is screwed. One side of the steel part is aligned with the socket, the other side overlaps. To test a piece of L-shaped insulating laminate it is inlayed to the L-shape block of steel. To hold down the insulation laminate a smaller plastic part is pressed by two screws against a plain piece of copper flat wire, that is placed on top of the paper. This creates the layered fixed structure: copper flat wire - insulation paper - mild steel. The described test rig is shown in Fig. 5. The final grooving result is mainly influenced by the following parameters: \u2022 Type of insulation material \u2022 Fiber orientation of aramid \u2022 Grooving width \u2022 Grooving depth \u2022 Geometry of grooving/die rollers Since the parameter range is large, this paper examines those which probably have the greatest influence on the grooving and its influence on the breakdown voltage. The grooving depth and grooving width are obviously the most significant parameters, as they actively adjust the degree to which the paper is deformed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003934_ecce44975.2020.9235932-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003934_ecce44975.2020.9235932-Figure3-1.png", + "caption": "Figure 3: Temperature estimation result.", + "texts": [ + " Meanwhile, the inductance estimation errors are both less than 0.5%, which also verifies the reliability. (6) (7) (8) (9) (10) (11) (12) (13) (14) 5823 Authorized licensed use limited to: Auckland University of Technology. Downloaded on December 23,2020 at 08:02:16 UTC from IEEE Xplore. Restrictions apply. IV. FINITE ELEMENT ANALYSIS OF PM TEMPERATURE DISTRIBUTION In order to analyze the distribution of PM temperature during the operation of the PMSM, a simulation model of the test motor is built by ANSYS shown in Fig.3. The mechanical structure and materials are completely consistent with the PMSM to be tested. For the PM part, the material is set to NdFeB, as well as the thermal conductivity is set to 8.95 W/m\u00b7C. As for the simulated test condition, the speed and the load are both set to the rated values, of which the speed is 3000r/min and the load condition reflected on the stator phasecurrent is 2.89A. In addition, the total simulation time of motor operating is 60 minutes. The simulation results show that the maximum temperature of the PM rises from 22 to 73" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003461_embc44109.2020.9175945-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003461_embc44109.2020.9175945-Figure5-1.png", + "caption": "Fig. 5. Relationship between three fingers and lever operation of the operating interface", + "texts": [ + " The polyacetal guide pin of the finger moves in the guide track, which controls the trajectory of the finger. Conversely, three fingers are closed by returning the connector to the initial position by two extension springs (spring constant 0.49 N/mm) and one compression spring (spring constant 4.1 N/mm). Fig. 4 shows the inside of the operating interface. When the lever head is pushed by a user, the lever rotates around a rotation axis, and then the plunger of the small-diameter syringe is pushed. Three fingers are opened by pushing the lever of the operating interface, as shown in Fig. 5. The operating interface is mounted to the user\u2019s upper arm on the affected side with a band, and the user operates the lever by pinching it between the upper arm and the side, as shown in Fig. 6. The operating interface can be worn without a harness, unlike the body-powered prosthetic hand. 4948 Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 20,2020 at 16:51:50 UTC from IEEE Xplore. Restrictions apply. Lever State of openingState of closing Tripod grasp Cylindrical grasp Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure78.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure78.10-1.png", + "caption": "Fig. 78.10 Max. shear stresses of a existing and b modified FH", + "texts": [], + "surrounding_texts": [ + "The test is carried out on each sample at five different locations with three reputations. These results show that the hardness of the material is not the same throughout the hammer. At the hard-facing welded part, the hardness is high and less at the tail part. Due to this, the Fibrizer hammer frequently breaks at the welded part. The welding heat input is important welding parameter, which effects on the structure and properties of the weld metal. At location-1, sample harness result shows that the hardness varied between 110 and 124 VHN. At the location-2, Fibrizer hammer hardness is between 146 and 151 VHN, and at the location-3, hardness is between 152 and 167 VHN. 78.3.4 Cost Comparison Between Existing and Modified Fibrizer Table 78.6 indicates the cost comparison between existing and modified Fibrizer. The difference in cost between them is $11,088.00. 78.3.5 Harmonic Analysis Results of Existing and Modified Fibrizer Hammers ANSYS finite element software is used for the simulation to performing harmonic analysis. From the simulation results, total deformation, direction deformation, von Mises stresses, maximum shear stresses, and maximum amplitude are evaluated for existing and modified Fibrizer hammers. 942 T. Mathewos et al. From Figs. 78.7, 78.8, 78.9, and 78.10 analysis, results are tabulated in Table 78.7. From the results, it is understood that the total deformation in the modified Fibrizer hammer is lesser than the total deformation of the existing hammer. This shows that the modified Fibrizer hammer is more reliable than the current one. As the result shows that the maximum equivalent (von Mises) stress of current FH is increased by double, the higher the stress, the more the material will be exposed to be broken. The maximum shear stress result in the current Fibrizer hammer is more than the modified Fibrizer hammer. The smaller the radius (r) and web (h) in stepped plate of the Fibrizer hammer, the higher will be the stress. In Table 78.7, analysis results show that there is less total and directional deformation in modified Fibrizer hammer. The equivalent (von Mises) stress and max. shear stress result also much less by half from the existing. Similar manner frequency responses are verified for other surfaces also and tabulated in Table 78.8. From Table 78.8 data, the output of harmonic response has been taken in the form of amplitude which can be understood as mean stress development in any engineering components. Modified FH has more amplitude initially and gradually decreased compared to current FH. The overall behavior modified FH is taking less deformations and stresses by which life span of FH definitely will improve. 78 Design Analysis and Modification of Sugarcane Fibrizer Hammer \u2026 943 944 T. Mathewos et al." + ] + }, + { + "image_filename": "designv11_80_0001182_s13344-019-0054-0-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001182_s13344-019-0054-0-Figure2-1.png", + "caption": "Fig. 2.\u00a0\u00a0\u00a0Fin\u00a0numbers\u00a0and\u00a0deflections.", + "texts": [ + "\u00a0The\u00a0acceleration\u00a0of\u00a0the\u00a0origin\u00a0of\u00a0the\u00a0body\u00a0frame\u00a0in\u00a0x,\u00a0y\u00a0and\u00a0z-axes\u00a0of\u00a0the\u00a0body\u00a0frame which\u00a0are\u00a0respectively\u00a0marked\u00a0by\u00a0 ,\u00a0 \u00a0and\u00a0 \u00a0are\u00a0equal\u00a0to ax = u\u0307\u2212 vr+wq ay = v\u0307\u2212wp+ur az = w\u0307\u2212uq+ vp (2) Hence,\u00a0Eq.\u00a0(1)\u00a0can\u00a0also\u00a0be\u00a0stated\u00a0as: m [ ax + zCG(pr+ q\u0307) ] = X m [ ay+ zCG(qr\u2212 p\u0307) ] = Y m [ az\u2212 zCG(p2+q2) ] = Z Ix p\u0307+mzCGay = K Iyq\u0307+ (Ix \u2212 Iz)rp+mzCGax = M Izr\u0307+ (Iy\u2212 Ix)pq = N (3) 2.1\u00a0\u00a0Roll\u00a0dynamics KIn\u00a0Eq.\u00a0(3),\u00a0 \u00a0which\u00a0is\u00a0the\u00a0rolling\u00a0moment\u00a0about\u00a0the\u00a0origin\u00a0is\u00a0equal\u00a0to K = K\u03b4a\u03b4a+Kp p+K p\u0307 p\u0307+Kmot\u2212mgyzCG, (4) \u03b4awhere\u00a0 ,\u00a0which\u00a0is\u00a0the\u00a0deflection\u00a0of\u00a0the\u00a0aileron,\u00a0is\u00a0defined as\u00a0(Fig.\u00a02): Kmot zCG zCG gy \u00a0is\u00a0the\u00a0reaction\u00a0torque\u00a0caused\u00a0by\u00a0the\u00a0rotation\u00a0of propulsive\u00a0motor.\u00a0As\u00a0stated\u00a0before,\u00a0 \u00a0is\u00a0the\u00a0distance between\u00a0the\u00a0center\u00a0of\u00a0gravity\u00a0and\u00a0the\u00a0center\u00a0of\u00a0buoyancy. Generally,\u00a0the\u00a0AUVs\u00a0are\u00a0designed\u00a0in\u00a0a\u00a0way\u00a0that\u00a0the\u00a0CG\u00a0is exactly\u00a0under\u00a0the\u00a0center\u00a0of\u00a0buoyancy\u00a0(Rentschler\u00a0et\u00a0al., 2003,\u00a02006;\u00a0Hong\u00a0et\u00a0al.,\u00a02013).\u00a0In\u00a0this\u00a0way,\u00a0the\u00a0righting\u00a0moment\u00a0of\u00a0buoyancy\u00a0provides\u00a0roll\u00a0and\u00a0pitch\u00a0stability.\u00a0Hence, \u00a0should\u00a0be\u00a0positive.\u00a0 ,\u00a0which\u00a0is\u00a0the\u00a0gravity\u00a0component in\u00a0y-axis\u00a0of\u00a0the\u00a0body\u00a0frame,\u00a0is\u00a0equal\u00a0to gy = gcos\u03b8 sin\u03c6", + "\u00a0(7)\u00a0is\u00a0linearized\u00a0as: I\u2032x p\u0307 = K\u03b4a\u03b4a+Kp p+K\u03c6\u03c6, (8) I\u2032x = Ix \u2212K p\u0307 K\u03c6 = \u2212mgzCG p = \u03c6\u0307 where\u00a0 \u00a0and\u00a0 .\u00a0Hence,\u00a0by\u00a0assuming \u00a0and\u00a0applying\u00a0a\u00a0Laplace\u00a0transform,\u00a0the\u00a0transfer\u00a0function\u00a0is\u00a0governed\u00a0as: \u03c6 \u03b4a = K\u03b4a I\u2032xs2\u2212Kps\u2212K\u03c6 . (9) 2.2\u00a0\u00a0Yaw\u00a0dynamics According\u00a0to\u00a0Eq.\u00a0(1)\u00a0and\u00a0making\u00a0some\u00a0simplifying\u00a0as- sumptions,\u00a0the\u00a0linear\u00a0equations\u00a0of\u00a0motion\u00a0in\u00a0the\u00a0steering plane\u00a0can\u00a0be\u00a0achieved\u00a0as\u00a0(Hong\u00a0et\u00a0al.,\u00a02013): m(v\u0307 +ur) = Yv\u0307v\u0307+Yr\u0307 r\u0307+Yvv+Yrr+Y\u03b4r\u03b4r; (10) Izr\u0307 = Nv\u0307v\u0307 +Nr\u0307 r\u0307+Nvv+Nrr+N\u03b4r\u03b4r, (11) \u03b4rwhere\u00a0 \u00a0is\u00a0the\u00a0deflection\u00a0of\u00a0rudder,\u00a0and\u00a0it\u00a0can\u00a0be\u00a0obtained as\u00a0(Fig.\u00a02): \u03b4r = 1 2 (\u03b43\u2212\u03b41). (12) Eqs.\u00a0(10)\u00a0and\u00a0(11)\u00a0can\u00a0be\u00a0represented\u00a0in\u00a0state\u00a0space\u00a0as: v\u0307 r\u0307 \u03c8\u0307 = \u2212a11 \u2212a12 0 \u2212a21 \u2212a22 0 0 1 0 [ v r \u03c8 ] + b1 b2 0 \u03b4r (13) where a11 = \u2212 Yr\u0307Nv (Iz\u2212Nr\u0307)(m\u2212Yv\u0307) + Yv (m\u2212Yv\u0307) 1\u2212 Yr\u0307Nv\u0307 (Iz\u2212Nr\u0307)(m\u2212Yv\u0307) ; (14) a12 = \u2212 Yr\u0307Nr (Iz\u2212Nr\u0307)(m\u2212Yv\u0307) + Yr \u2212mu (m\u2212Yv\u0307) 1\u2212 Yr\u0307Nv\u0307 (Iz\u2212Nr\u0307)(m\u2212Yv\u0307) ; (15) a21 = Nv\u0307a11\u2212Nv Iz\u2212Nr\u0307 ; (16) a22 = Nv\u0307a12\u2212Nr Iz\u2212Nr\u0307 ; (17) b1 = Yr\u0307N\u03b4r (Iz\u2212Nr\u0307)(m\u2212Yv\u0307) + Y\u03b4r (m\u2212Yv\u0307) 1\u2212 Yr\u0307Nv\u0307 (Iz\u2212Nr\u0307)(m\u2212Yv\u0307) ; (18) b2 = Nv\u0307b1+N\u03b4r Iz\u2212Nr\u0307 . (19) Hence\u00a0by\u00a0applying\u00a0a\u00a0Laplace\u00a0transform,\u00a0the\u00a0transfer function\u00a0of\u00a0yaw\u00a0dynamics\u00a0can\u00a0be\u00a0governed\u00a0as: \u03c8 \u03b4r = b2s+ (b2a11\u2212a21b1) s3+ (a11+a22)s2+ (a11a22\u2212a21a12)s . (20) 2.3\u00a0\u00a0Pitch\u00a0dynamics According\u00a0to\u00a0Eq.\u00a0(1),\u00a0and\u00a0making\u00a0some\u00a0simplifying\u00a0assumptions,\u00a0the\u00a0linear\u00a0equations\u00a0of\u00a0motion\u00a0in\u00a0the\u00a0diving plane\u00a0can\u00a0be\u00a0achieved\u00a0as\u00a0(Hong\u00a0et\u00a0al.,\u00a02013): m(w\u0307 \u2212uq) = Zw\u0307w\u0307+Zq\u0307q\u0307+Zww+Zqq+Z\u03b4e\u03b4e; (21) Iyq\u0307 = Mw\u0307w\u0307+Mq\u0307q\u0307+Mww+Mqq+M\u03b4e\u03b4e\u2212mgzCG \u03b8; (22) z\u0307 = w\u2212u0\u03b8, (23) \u03b4ewhere\u00a0 \u00a0is\u00a0the\u00a0deflection\u00a0of\u00a0elevator,\u00a0and\u00a0is\u00a0obtained\u00a0as (Fig.\u00a02): \u03b4e = 1 2 (\u03b42\u2212\u03b44) (24) u0and\u00a0 \u00a0is\u00a0the\u00a0steady-state\u00a0speed\u00a0of\u00a0the\u00a0AUV.\u00a0Eqs.\u00a0(21),\u00a0(22) and\u00a0(23)\u00a0can\u00a0be\u00a0represented\u00a0in\u00a0state\u00a0space\u00a0as: w\u0307 q\u0307 \u03b8\u0307 z\u0307 = \u2212c11 \u2212c12 \u2212c13 0 \u2212c21 \u2212c22 \u2212c23 0 0 1 0 0 1 0 \u2212u0 0 w q \u03b8 z + d1 d2 0 0 \u03b4e (25) c11 = \u2212 Zq\u0307Mw (Iy\u2212Mq\u0307)(m\u2212Zw\u0307) + Zw (m\u2212Zw\u0307) 1\u2212 Zq\u0307Mw\u0307 (Iy\u2212Mq\u0307)(m\u2212Zw\u0307) ; (26) c12 = \u2212 Zq\u0307Mq (Iy\u2212Mq\u0307)(m\u2212Zw\u0307) + mu+Zq (m\u2212Zw\u0307) 1\u2212 Zq\u0307Mw\u0307 (Iy\u2212Mq\u0307)(m\u2212Zw\u0307) ; (27) c13 = Zq\u0307mg\u03beZ (Iy\u2212Mq\u0307)(m\u2212Zw\u0307) 1\u2212 Zq\u0307Mw\u0307 (Iy\u2212Mq\u0307)(m\u2212Zw\u0307) ; (28) c21 = Mw\u0307c11\u2212Mw Iy\u2212Mq\u0307 ; (29) c22 = Mw\u0307c12\u2212Mq Iy\u2212Mq\u0307 ; (30) c23 = Mw\u0307c13+mg\u03beZ Iy\u2212Mq\u0307 ; (31) d1 = Zq\u0307M\u03b4e (Iy\u2212Mq\u0307)(m\u2212Zw\u0307) + Z\u03b4e (m\u2212Zw\u0307) 1\u2212 Zq\u0307Mw\u0307 (Iy\u2212Mq\u0307)(m\u2212Zw\u0307) ; (32) d2 = Mw\u0307d1+M\u03b4e Iy\u2212Mq\u0307 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001552_s12555-019-0234-y-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001552_s12555-019-0234-y-Figure5-1.png", + "caption": "Fig. 5. Schematic diagram of a 2-DOF discrete bending joint composed of 2N+1 rolling unit joints.", + "texts": [ + " Notably, the proposed method using 2N + 1 units guarantees the best continuum-like bending performance with a completely alternate stacking sequence of Y-P-. . . -P-Y or P-Y-. . . -P-Y without requiring optimization of the stacking sequence. This method can be considered an easy application of the basic stacking scheme discussed in the Introduction: Y-P -. . . - Y-P [16\u201321]. Because the proximal and distal unit joints have yawing motion, the proposed stacking sequence is Y-P. . . -P-Y. 3.1. 2-DOF discrete bending model using 2N+1 units A discrete bending joint composed of 2N + 1 rolling units is illustrated in Fig. 5, where odd- and evennumbered units have yawing and pitching motions, respectively. The coordinate frames of the intermediate linkages are located at the center of each linkage for easy com- prehension. The kinematics of this joint can be expressed in terms of the modified Denavit\u2013Hartenberg (D\u2013H) parameters listed in Table 2 by using the Craig method [25]. Notably, for intuitive and reasonable coordinate generation, the coordinate frames of the proximal and distal units in a straight posture should be oriented along the same direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001068_acc.2019.8815125-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001068_acc.2019.8815125-Figure1-1.png", + "caption": "Fig. 1. Boat diagram and actuator detail.", + "texts": [], + "surrounding_texts": [ + "Consider a fixed inertial frame {I} and a body frame {B} attached to the boat\u2019s center of mass. The configuration of {B} with respect to {I} is given by the pair (R,p) = ( I BR , IpB). to the x\u2212 y plane. The trajectory tracking objective is to drive a fixed point in the boat body frame, \u03b4, to follow a desired reference trajectory pd(t) : R \u2192 R2. The available actuations are the force provided by each motor propeller and the motor angle \u03b8. The orientation of the vehicle during the tracking maneuver is not a priori prescribed. Since the boat is underactuated, e.g. is not possible to generate a force with only lateral component, its final orientation will depend on the dynamics of the closed-loop system and the trajectory to be followed." + ] + }, + { + "image_filename": "designv11_80_0000790_iraniancee.2019.8786525-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000790_iraniancee.2019.8786525-Figure3-1.png", + "caption": "Fig. 3. Mechanical power characteristics of a wind turbine", + "texts": [ + " The three main reasons for having an active pitch control system are: a) Optimizing the output power of the system b) Adjusting all system variables at their limits and stability c) Minimizing the load effect on the mechanical components Pitch control limits the output power when the wind speed is over the rated value. The electrical power produced in the wind turbines are related to the cube of the wind speed. The power coefficient or shows the extractable part of the wind power and is a function of tip speed ratio and pitch angle. In fig. 3 the mechanical power characteristics of a wind turbine is shown as a function of rotor and wind speed [12]. The operating regions of the wind turbine system is shown in fig. 4. When wind speed increases and becomes more than the rated value, the pitch control system adjust the rotor speed at its rated value by increasing the pitch angle [12]. If the power coefficient of wind turbine is very low, frequency changes according to wind speed changes, are negligible. But if the power coefficient of wind turbine is high, frequency changes according to wind speed changes will increase [13]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000113_robio.2018.8665227-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000113_robio.2018.8665227-Figure1-1.png", + "caption": "Fig. 1. Schematic of our quadrotor", + "texts": [ + " In Section II, the simplified mathematical modeling of a quadrotor is formulated. In Section III, modeling errors are introduced and a KF-based active modeling method is proposed for a QSL system. In Section IV, experimental data is used to identify unknown parameters and validate the remarkable performance of the proposed modeling technique considering the modeling errors. In Section V, some concluding remarks, together with the future work, are briefly given. The studied QSL system is based on the developed quadrotor with a cross \u2018X\u2019 configuration as shown in Fig. 1. To build the mathematical model of the quadrotor, two frames are introduced, including the body-fixed reference frame and the world-fixed inertial reference frame. In this paper, we model the quadrotor by utilizing Newton-Euler method from the steering motor input to the position and attitude. Based on our previous work [3] and relate references [13], [15], 978-1-7281-0377-8/18/$31.00 \u00a9 2018 IEEE 2279 the dynamic equation of the quadrotor is described by the following differential equations:\u23a7\u23a8 \u23a9 R\u0307 = RS (\u03d1) J\u03d1\u0307+ S (\u03d1)\u03d1 = M\u03c4 MP\u0308 = RF1 +MG ", + " That is, a practical KF-based active modeling method for a QSL system in three dimensional situations is proposed on the basis of the modeling of a quadrotor and parameters identification. In this section, we will identify unknown model parameters, validate the existence of modeling errors and verify the effectiveness of the proposed active modeling method on a QSL experimental platform. In this subsection, we will identify related parameters in Eq. (6) and use experimental data to verify the identification results. Shown in Fig. 1, the experimental setup we use include a quadrotor with cross \u2018X\u2019 configuration, a Pixhawk flight controller and the Odroid-XU4. By using the quadrotor, two groups of hovering flights are conducted, in which the hovering point is set as (0, 0, 0.8). Then, one group of data is utilized for system identification, and the other group is for validation. Least square method is used to carry out parameter identification [18], [19]. And we can get M = 1.484kg, g = 9.8m/s2, l = 0.225m, Jxx = 0.0042948kg \u00b7 m2, Jyy = 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000591_978-3-030-17763-8_17-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000591_978-3-030-17763-8_17-Figure3-1.png", + "caption": "Fig. 3 The vehicle for experiment and its setup", + "texts": [ + " The vehicle we use is a quadcopter similar to DJI FlameWheel 450. The diameter of quadcopter is approximately 45 cm and it is a suitable size for indoor flight experiment. The load capacity of our vehicle is approximately 2.5 kg, which makes it possible to equip quadcopter with other peripheral devices (like camera, etc.). As for endurance, the flight time of a single flight can be up to more than 10 min, which is enough for most of the experiments. Some peripheral devices are attached to the vehicle, depicted in Fig. 3. The flight controller we choose is the Pixhawk (depicted in Fig. 4a), which is produced by the famous open-hardware manufacturer 3DR. Pixhawk is an industry standard autopilot and has a stable and high-level calculation performance. In addition, Pixhawk owns a variety of input and output interfaces, which provides a great convenience for equipping quadcopter with other peripheral devices. The onboard processor we choose is Odroid-XU4, which is a heterogeneous multi-processing (HMP) octa core linux computer (depicted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002906_pen.25409-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002906_pen.25409-Figure12-1.png", + "caption": "FIGURE 12 Predicted parison shape at time = 160 seconds, Weissenberg number Wi = 40 and for a die exit cross-sectional aspect ratio \u03c7 = 62.5, using a viscoelastic PTT model, A, extrudate shape, B, cross-sectional shape of the extruded sheet at a distance of 60 mm from the die exit showing sinusoidal wrinkle deformation of large amplitude", + "texts": [ + "0565 MPa is found just after the die exit, while the critical compressive stress for the sheet to wrinkle is \u03c3c = \u22120.0524 MPa (from energy method). As the compressive stress induced in the sheet next to the die exit is slightly below the critical compressive stress for the extruded sheet to wrinkle, that is, N1 < \u03c3c, wrinkles of very low amplitude occur, as highlighted in Figure 11. By increasing further the Weissenberg number Wi to 40 (corresponding to Vmax = 5.5 mm/s in Equation (39)), sinusoidal wrinkles of high amplitude develop soon after the sheet is extruded from the die exit as shown in Figure 12. The distribution of the first normal stress difference N1 along the central line, in the flow axis, is presented in Figure 9D. It starts from zero in the die, increases to reach a maximum just before the die exit, then decreases to become negative and reach a minimum negative value just after the die exit. After a certain distance from the die exit, the normal and shear stresses relax to zero in absence of confinements. Therefore, soon after extrusion from the die exit, the extruded sheet behaves as a rectangular sheet subjected to a compressive residual stress in the width direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001924_icoecs46375.2019.8949914-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001924_icoecs46375.2019.8949914-Figure1-1.png", + "caption": "Fig. 1. The topology of a 9-phase bearingless EM", + "texts": [ + " However, the development of control technology and power electronics allowed the creation of multi-phase EMs in which one winding is used to control two decoupled d-q planes and allows to ensure the fault tolerance by one winding. In particular, a bearingless 5-phase DC motor is presented in [32], where the multiphase winding is used to control two decoupled d-q planes and creates torque and radial force, respectively. In [33], a 9-phase bearingless EM was developed in a similar way. The topology of this EM is shown in Fig. 1. Thus, it can be concluded that the technology of nonbearing electrical machines, as well as the technologies of fault-tolerance AMBs can solve the problems of failures in the EM bearings. At the same time, the development of diagnosis of mechanical bearing supports can also minimize failures in EM. The main failures on the stator are the failure of one of the phases or its exclusion from the energy conversion process, as well as short circuits at the terminals of the EM winding or in the EM control system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003838_s11071-020-06016-4-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003838_s11071-020-06016-4-Figure6-1.png", + "caption": "Fig. 6 Disk in a half-pipe", + "texts": [ + " The part of it right up to the intersection with line 1 corresponds to the triple dynamics: generally speaking rolling can continue or terminate at any point of this curve section. Generally speaking, there is a continuum solutions to equations of motion. At the same time, the straightforward approach missed detecting them and leads to the single (of continuum) solution: rolling is followed by sliding (at line 1) with subsequent jump. A further modification of the problem is a curvilinear support. We consider a half-pipe of radius R (Fig. 6) as the simplest example of such type. There exist a family of periodic motions where the disk preserves rolling. Each of them depends on two parameters: total energy and the angular position of the disk h0 in the lowest position of his geometric center C. We will check whether these motions are really possible without violation of the contact conditions. Now let h be the angle between CG and the tangent direction at the point of contact, u be the angle between vertical and line of centers. The absolute angular velocity of the disk is x \u00bc _u\u00fe _h, and the condition of rolling is R _u\u00fe r _h 0 ) u \u00bc k h h0\u00f0 \u00de; k \u00bc r=R\\1 \u00f05:1\u00de Equations of motion can be derived by analogy with Sect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002050_iros40897.2019.8968099-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002050_iros40897.2019.8968099-Figure4-1.png", + "caption": "Fig. 4. (a) a new configuration with tilted rotors. (b) the positive direction of the tilting angle \u03b1.", + "texts": [ + " Instead, one can yaw by tilting the rotors by an angle \u03b1 about the x-axis of their corresponding motor frame, thus, using a small component of the propeller\u2019s thrust force to generate relatively larger yaw moments. Note that the tilting angle should be small enough (\u22120.2 < \u03b1 < 0.2 rad) so that (8) and (9) still hold true and the component of the thrust force that balances the weight of the vehicle experiences relatively small changes. We introduce a tilting angle \u03b1i about the x-axis of the motor frame Mi to the rotors similar to that in [3] as shown in Fig. 4. According to [3], having \u03b11,3 > 0 and \u03b12,4 < 0 adds to the passive stability of the vehicle in yaw motion which is an improvement in quadcopter flight without any rotor failure. But we are interested in finding the effects of this tilting angle on the mechanical power of the quadcopter after rotor failure and also on spinning UAVs like bispinner. A new configuration is proposed by tilting the rotors about the x-axis of the motor frame (shown in blue in Fig. 1) as shown in Fig. 4 (a) where the positive direction of the tilting angle \u03b1i is shown in Fig. 4 (b). Because rotors 1 and 3 are assumed to be turning in the negative direction of z-axis of the body frame, by tilting these motors by any positive angle, the vehicle tends to generate a yaw motion that is in favor of reducing mechanical power (11). Whereas for rotors 2 and 4 which are turning in the positive direction of the z-axis of the body frame, the tilting angle should be negative. Note that, for simplicity, it is assumed |\u03b11| = |\u03b12| = |\u03b13| = |\u03b14|. Although this new configuration might not lead to reducing mechanical power in quadcopters without rotor failure very effectively, however, it can be drastically useful in quadcopters experiencing rotor failure and in spinning UAVs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure81.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure81.2-1.png", + "caption": "Fig. 81.2 Blood warmer assembly", + "texts": [ + " Humidity effects were not considered in the simulation. The voltage difference was applied on the coil path of the top plate and the bottom plate (to generate joule heating in the plates). A perfect thermal contact was assumed between the disposable and heater plates. The blood warmer assembly used in the dialysis machine consists of four components, the fluid domain, blood flow path, the heater coil and the two aluminium heater plates. The material properties of the different components used in the simulations are as listed in Table 81.1 (Fig. 81.2). The power required for the flow rates 100, 150, 200, 250, 300, 350, 400, 450 (ml/ min) are 60, 89, 118, 148, 177, 206, 234, 264 (W), respectively, as listed in the Table 81.2. 81 Design of Medical Device Product Using Multiphysics Simulations 969 970 R. Kapuganti There is a linear directly proportional relationship between the mass flow rate and heat energy required as depicted in the Fig. 81.3. The spatial thermal distribution of the blood flow path is depicted in the Fig. 81.4. The temperature of the fluid increases along the blood flow passage" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003185_012018-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003185_012018-Figure5-1.png", + "caption": "Figure 5. Magnetic field distribution for a four-spring harvester with two by two spaced magnets at zero horizontal deviation.", + "texts": [ + " The studied harvesters are nonlinear mechanical oscillating systems. Their simulations made by ANSYS R19.1 take into account the fixture, the gravity effect of the plates with permanent magnets and the mechanical characteristics of the used springs. The horizontal deviation x of the mechanical system \u201cmass (plate with permanent magnets) \u2013 springs\u201d was obtained while modeling the four-spring electromagnetic harvesters, Figure 4. The magnetic field distribution of the three electromagnetic harvesters was obtained by means of FEMM 4.2. Figure 5 shows the magnetic field distribution for the third four-spring harvester with two spaced magnets in two places at zero horizontal deflection, and Figure 7 presents the distribution at maximum deviation. Figure 6 and figure 8 illustrate the magnetic flux density changes along the length of the harvester coil at zero and maximum horizontal deviation. From Figure 6 it can be seen that at zero horizontal deviation the normal magnetic flux density is zero, and at maximum deviation the maximum magnetic flux density Bmax is obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002614_j.promfg.2020.04.235-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002614_j.promfg.2020.04.235-Figure7-1.png", + "caption": "Fig. 7. (a) Bevel gear after the numerical simulation of the forming process, (b) section representation of (a), (c) plastic strain distribution that results after the numerical simulation of the forming process", + "texts": [ + " Computed coefficients for the used flow curve approach HenselSpittel-10 for the materials 41Cr4 conventional, 41Cr4 deposition and X45CrSi9-3 deposition 41Cr4 conventional 41Cr4 deposition X45CrSi9-3 deposition A [Ns/mm2] 467044 0.0189 0.0250 m1 [1/\u00b0C] -0.0026 -0.0059 -0.0063 m2 [-] 0.1859 0.0385 0.2973 m3 [-] -0.2563 0.0001 0.0010 m4 [-] -0.0110 -0.0058 0.0049 m5 [1/\u00b0C] -7.42\u2219 1 5 0.0012 0.0006 m6 [-] -0.8706 -1.456 -1.3861 m7 [1/\u00b0C] 0.0003 0.0001 0.0001 m8 [-] -0.7103 2.2025 2.3159 Fig. 6. Force-displacement curve of the uniaxial compression test In Fig. 7 (a) the bevel gear after the forming process is demonstrated, which was achieved with an initial inner layer thickness of 1 mm and an initial outer layer thickness of 3 mm. In the area of the formed tooth, only the materials of the welded layers are found. Fig. 7 (b) shows that the radial material flow T = 1050 C 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 50 0 100 150 200 250 300 350 Plastic strain \ud835\udf11\ud835\udf11 Fl ow st re ss in M Pa X45CrSi9-3 ?\u0307?\ud835\udf11 = 1 T = 900 C T = 1200 C T = 900 C T = 1200 C T = 1050 C 0 2 4 6 8 0 1 2 3 4 5 Fo rc e in k N Displacement in mm Experiment / Simulation X45CrSi93-3 deposition welding T = 900 C ?\u0307?\ud835\udf11 = 1 T = 900 C ?\u0307?\ud835\udf11 = .1 T = 1200 C ?\u0307?\ud835\udf11 = .1 Experiment / Simulation 41Cr4 conventional welding Experiment / Simulation 41Cr4 deposition welding/ / / Fo rc e in k N ", + " Experiment / Simulation X45CrSi9-3 depos tion welding material Experiment / Simulation 41Cr4 conventional ro material Experiment / Simulation 41Cr4 deposition welding material ?\u0307?\ud835\udf11 = 1 Fl ow st re ss \ud835\udc58\ud835\udc58 \ud835\udc53\ud835\udc53 in M Pa X45CrSi9-3 depends on the position in z-axis direction. In the lower area, the material is pressed strongly in the radial direction. The conical shape of the bevel gear and the selected conical shape of the bevel gear base caused this material distribution after the forming process. The strongest radial material flow occurs in the area of the burr. The distribution of the plastic strain is shown in Fig. 7 (c). A degree of plastic strain of up to 5.06 is achieved in the area of the burr. In Fig. 9, the results of the numerical investigation with the initial layer thickness of the inner and outer layer according to Table 1 are presented. Fig. 9 shows a sectional view of the numerical results, which shall help to better pin the resulting material distribution in the tooth of the bevel gear after the forming process. The aim of the numerical investigation was the analysis of the material distribution after the forming process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003853_ecai50035.2020.9223138-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003853_ecai50035.2020.9223138-Figure11-1.png", + "caption": "Fig. 11. Chosen air fan \u2013 Sunon MAglev MEC0381V1", + "texts": [ + " The ventilator chosen to supply the decontaminated air should normally allow about 1/2 liter of air to enter the respirator during respiration and is eliminated in the same amount. Considering that a healthy adult breathes 16 times a minute, it can be calculated that in an hour about 480 liters of air are carried through the lungs. From this air, an adult consumes about 300 cm3 of oxygen per minute, which is 18 liters of oxygen in an hour. The chosen fan, model Sunon MAglev MEC0381V1, 120mm, shown in the figure 11 offers an air flow between 12 and 75 cubic meters / minute . This is equivalent to an air flow of 720-4500 liter/hour. Fabricantul specifica o durata de functionare de 70.000 ore. The power consumption of the fan is 10W / The electrical efficiency of the converter of a dial used in the control of the fan speed is about 98%. DC motor speed variator To change the fan speed and implicitly to change the air flow we use a DC-DC converter [15-19]. The electrical characteristics of this converter are the following: - supply voltage: 4" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure78.7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure78.7-1.png", + "caption": "Fig. 78.7 Total deformation of a existing and b modified FH", + "texts": [], + "surrounding_texts": [ + "The test is carried out on each sample at five different locations with three reputations. These results show that the hardness of the material is not the same throughout the hammer. At the hard-facing welded part, the hardness is high and less at the tail part. Due to this, the Fibrizer hammer frequently breaks at the welded part. The welding heat input is important welding parameter, which effects on the structure and properties of the weld metal. At location-1, sample harness result shows that the hardness varied between 110 and 124 VHN. At the location-2, Fibrizer hammer hardness is between 146 and 151 VHN, and at the location-3, hardness is between 152 and 167 VHN. 78.3.4 Cost Comparison Between Existing and Modified Fibrizer Table 78.6 indicates the cost comparison between existing and modified Fibrizer. The difference in cost between them is $11,088.00. 78.3.5 Harmonic Analysis Results of Existing and Modified Fibrizer Hammers ANSYS finite element software is used for the simulation to performing harmonic analysis. From the simulation results, total deformation, direction deformation, von Mises stresses, maximum shear stresses, and maximum amplitude are evaluated for existing and modified Fibrizer hammers. 942 T. Mathewos et al. From Figs. 78.7, 78.8, 78.9, and 78.10 analysis, results are tabulated in Table 78.7. From the results, it is understood that the total deformation in the modified Fibrizer hammer is lesser than the total deformation of the existing hammer. This shows that the modified Fibrizer hammer is more reliable than the current one. As the result shows that the maximum equivalent (von Mises) stress of current FH is increased by double, the higher the stress, the more the material will be exposed to be broken. The maximum shear stress result in the current Fibrizer hammer is more than the modified Fibrizer hammer. The smaller the radius (r) and web (h) in stepped plate of the Fibrizer hammer, the higher will be the stress. In Table 78.7, analysis results show that there is less total and directional deformation in modified Fibrizer hammer. The equivalent (von Mises) stress and max. shear stress result also much less by half from the existing. Similar manner frequency responses are verified for other surfaces also and tabulated in Table 78.8. From Table 78.8 data, the output of harmonic response has been taken in the form of amplitude which can be understood as mean stress development in any engineering components. Modified FH has more amplitude initially and gradually decreased compared to current FH. The overall behavior modified FH is taking less deformations and stresses by which life span of FH definitely will improve. 78 Design Analysis and Modification of Sugarcane Fibrizer Hammer \u2026 943 944 T. Mathewos et al." + ] + }, + { + "image_filename": "designv11_80_0001233_systol.2019.8864780-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001233_systol.2019.8864780-Figure1-1.png", + "caption": "Figure 1. 6 DOF AUV angular and translational motions [2].", + "texts": [ + " Restrictions apply. control effectiveness factors (for example, partial loss of a control surface - when a part of the control surface breaks off - deformation of the control surface, etc.). AUV modeling is fairly complicated, and an exact analysis is only possible by including the underlying infinite dimensional dynamics of the surrounding fluid (sea water). AUVs move in 6 degrees of freedom (6DOF) since six independent coordinates are necessary to determine the position and orientation of a rigid body (See Figure 1) [2]. The first three coordinates and their time derivatives are of translational motion along the x, y and z-axes, while the last three coordinates ( , , ) and time derivatives are used to describe orientation and rotational motion. The linearized model of torpedo will be used instead of sample AUV in calculations. 6 different motion variables help to determine position and orientation. First three coordinates (x, y, z) are used to determine the position. Time derivatives of three coordinates (u, v, w) define transitions along x, y and z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003824_aiea51086.2020.00073-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003824_aiea51086.2020.00073-Figure1-1.png", + "caption": "Figure 1 Working principle diagram of J-type clamp 1-bolt; 2, 3-J component; 4-branch cable; 5-nut; 6-washer; 7-main cable", + "texts": [ + " The branch cable installation clamping tool proposed in this paper is mainly used for the live installation of various types of J-type clamps. Therefore, this article takes J-type clamps as an example for research. J-type clamp is a bolt-clamped permanent elastic clamp. Its unique structure and angle design ensure that the clamp has long-lasting clamping force, strong conductivity, and has outstanding characteristics such as controllable installation quality and visual installation[5]. As shown in Figure 1, the Jshaped clamp consists of two wedge-shaped J elements, bolts, nuts and gaskets. The J-type clamp uses a wedge-shaped constant contact pressure technology. The preload of the bolt and the force of the two J elements acting laterally on the wire do not interfere with each other. Even if the spring gasket of the bolt gradually loses its elasticity during operation, the tightening force is It will be reduced, but the contact performance will not be significantly affected, and it is less 321 978-1-7281-8288-9/20/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002099_icar46387.2019.8981581-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002099_icar46387.2019.8981581-Figure2-1.png", + "caption": "Fig. 2: Comparison between (a): an Admittance Control implemented at the robot\u2019s TCP(Sxyz) and (b): a Multi-Surface Admittance control", + "texts": [ + " These uncertainties, mixed with complex shape handled parts, and large-scale objects, make multi-surface contact tasks hard to achieve. In this regard, a new strategy to control the pose (position and orientation) of the robot\u2019s end-effector is proposed, taking into account the exerted forces on each surface of the handled part. This property allows to model the contact at each surface with a second-order mass-springdamper system. The parameters of this system are defined according to the environment stiffness and the desired contact dynamics. Summarizing, one admittance control schema is created for each of these surfaces. Figure 2 shows the comparison between an admittance control approach implemented on the TCP (Sxyz) of a robot (2a), and our Multi-Surface admittance control (MSAC) approach (2b) which allows to implement n admittance controller systems. The former calculates the object\u2019s position/velocity using one set of stiffness, damping an inertia constants at one point, which is well adapted for small handled objects. By contrast, our approach will employ different sets of constants, one set at each contact surface, to achieve the desired contact among the prominent and complex-shaped part and the environment", + " In this way, the coordinate vector representing the spatial force is given by: fO = ( nO f ) (4) and the linear/angular velocity by the spatial velocity: x\u0307O = ( \u03c9 \u03c5O ) (5) Considering a Cartesian coordinate frame Oxyz, hence f = ( fx fy fz )T is the force acting on a line passing through O\u2019s origin, nO = ( nxO nyO nzO )T is the moment of the force system about O, \u03c9 = ( \u03c9x \u03c9y \u03c9z )T is the body\u2019s angular velocity, and \u03c5O = ( \u03c5xO \u03c5yO \u03c5zO )T is the body\u2019s linear velocity of the body-fixed point at O\u2019s origin. As this control schema is based on an admittance controller, it is needed to calculate the desired spatial force acting on the concerned surface (fdU ) to achieve the interaction task. Take the robot\u2019s handled object exposed in Figure 2b as an example. Consider the goal of the interaction task to have a surface-surface contact between this object and the environment. At the end of a successful interaction task, the spatial force (fdS ) measured by a force sensor in the frame Sxyz, would be the desired spatial force applied to the rigid-body. Therefore, to calculate the desired spatial force acting at each surface, fdS is transformed into the surface\u2019s coordinate system (Uxyz). Hence, for the ith surface, the desired spatial force can be computed like: fdUi =Ui T f S \u2217 fdS (6) where UiT f S is a 6x6 transformation matrix for spatial forces from a frame Sxyz, to a frame Uixyz [14]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000355_iicpe.2018.8709462-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000355_iicpe.2018.8709462-Figure1-1.png", + "caption": "Fig. 1. Rolling-element bearing structure and parameters.", + "texts": [ + " Here, stator currents are 0i s qd ; rotor currents are 0 r qdi ; rotor speed r\u03c9 ; load torque lT ; and input stator voltage 0.s qdv Equation (1) represents the normal behavior of the induction motor, which is used in Extended Kalman filter formulation for state estimation. III. VALID BEARING FAULT MODELS A. Bearing Fault Types and their Characterstics Frequencies As most of the induction motors use rolling-element bearings. The schematic representation of rolling-element bearings structure is shown in Fig. 1. The main rollingelement bearing parameters are outer raceways, inner raceways, the balls, and the cage. Bearing faults can be classified into two types: 1) single-point faults, and 2) generalized roughness (permanently damaged bearing) [5]- [9]. Further, single-point bearing defects can be classified as follows: \u2022 Cage bearing fault \u2022 Outer raceway bearing fault \u2022 Inner raceway bearing fault \u2022 Ball bearing fault Each type of bearing fault introduces characteristics frequency ( )cf components in the stator current spectrum" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002334_j.matpr.2020.02.819-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002334_j.matpr.2020.02.819-Figure1-1.png", + "caption": "Fig. 1. FEA model of freight wagon intende", + "texts": [ + " wagon Freight wagon intended for transport of petroleum products is modelled using the FEMAP software with NX Nastran solver [6]. According to the construction type, shell elements of the appropriate thickness, 3D elements (for modelling of support plate, compensating ring, traction stop) and beam elements for modelling bolted connections were used for creating the finite element mesh. Structure is modelled in details with 236,800 elements and 243,846 nodes. General element side length is about 30 mm. The MKE model includes 14 contact pairs and 68 bolted connections. Fig. 1 shows the 3D FEA model of the whole wagon without bogies. Colors in Fig. 1 match the various thicknesses of shell elements. Because of structure symmetry a half of the model was used almost for all load cases, taking in consideration correspondent symmetry of the load cases. Full model was used for unsymmetrical load cases. The wagon structure is made so that there are three basic connection between its parts: connection between the tank and saddle (contact pair), middle connection (combination \u2013 contact pair and sliding bolt connection) and side connections between chassis of wagon and tank (combination \u2013 contact pair and sliding bolt connection; combination \u2013 contact pair and fixed bolt connection on Please cite this article as: M" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002648_j.mechmachtheory.2020.103891-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002648_j.mechmachtheory.2020.103891-Figure3-1.png", + "caption": "Fig. 3. Adaptation of mechanism of the second one-rotational and two-translations DOFs asymmetrical PKM family with mobility F N = 3 and C N = 3 , originally introduced in [38] (Fig. 2).", + "texts": [ + " In [10] , Qin and Dai investigated the configuration and the actuation of the 2US+UPS asymmetrical parallel mechanism, proposing two actuation schemes. In general, in overconstrained asymmetrical parallel mechanisms not all possible sets of joints can be selected as a valid actuation scheme. Furthermore, not always the actuators near the base are a valid actuation scheme which satisfies Condi- tion 1 . Refaat et al. in [38] introduce four families of three degrees of freedom translational-rotational asymmetrical parallel mechanisms. The mechanism presented in Fig. 3 was originally introduced by Refaat et al. in [38] ( Fig. 2 in the original publication). This mechanism has three degrees of freedom and three redundant constraints. Intuitively, if the revolute joints a and i and the prismatic joint e were actuated and locked, the platform 4 is still capable of rotating along the y axis. Thus, the natural set { a, e, i } is not a valid actuation scheme. On the other hand, if the joints { a, i } and one revolute joints in the set { f, g, h } were actuated, the mechanism is frozen, i", + " Set # { j, k, l } 1 { b, k, f } 26 { j, d, h } 51 { i, b, g } 76 { a, j, l } 2 { a, j, f } 27 { k, d, h } 52 { i, c, g } 77 { a, k, l } 3 { a, k, f } 28 { j, c, h } 53 { i, d, g } 78 { j, l, b } 4 { j, l, f } 29 { k, c, h } 54 { i, e, g } 79 { b, k, l } 5 { k, l, f } 30 { j, b, h } 55 { i, j, f } 80 { j, c, l } 6 { j, k, f } 31 { b, k, h } 56 { i, k, f } 81 { k, l, c } 7 { j, e, g } 32 { a, j, h } 57 { i, l, f } 82 { j, c, d } 8 { k, e, g } 33 { a, k, h } 58 { a, i, f } 83 { k, d, c } 9 { j, d, g } 34 { j, l, h } 59 { i, b, f } 84 { j, d, b } 10 { k, d, g } 35 { k, l, h } 60 { i, c, f } 85 { b, k, d } 11 { j, c, g } 36 { j, k, h } 61 { i, d, f } 86 { a, j, d } 12 { k, c, g } 37 { i, j, h } 62 { i, e, f } 87 { a, k, d } 13 { j, g, b } 38 { i, k, h } 63 { i, l, e } 88 { j, k, d } 14 { b, k, g } 39 { i, l, h } 64 { i, d, e } 89 { j, d, e } 15 { a, j, g } 40 { a, i, h } 65 { i, j, d } 90 { k, d, e } 16 { a, k, g } 41 { i, b, h } 66 { i, k, d } 91 { j, l, e } 17 { j, l, g } 42 { i, c, h } 67 { a, i, d } 92 { k, l, e } 18 { k, l, g } 43 { i, d, h } 68 { i, b, d } 93 { j, e, f } 19 { j, k, g } 44 { i, e, h } 69 { i, c, d } 94 { k, e, f } 20 { j, g, h } 45 { i, f, h } 70 { i, c, l } 95 { j, d, f } 21 { k, g, h } 46 { i, g, h } 71 { i, b, l } 96 { k, d, f } 22 { j, f, h } 47 { i, j, g } 72 { a, i, l } 97 { j, c, f } 23 { k, f, h } 48 { i, k, g } 73 { i, j, l } 98 { k, c, f } 24 { j, e, h } 49 { i, l, g } 74 { i, k, l } 99 { j, f, b } 25 { k, e, h } 50 { a, i, g } 75 All valid actuation schemes for the mechanism in Fig. 3 can be enumerated by applying Algorithm 1 . For this purpose, a non-singular configuration of the PKM family mechanism is considered. Regard that the enumeration of all actuation valid schemes is independent from the configuration selected, whenever only non-singular configuration are analyzed. Table 2 presents the coordinates of each joints for the configuration considered. Matrices [ \u02c6 M D ] 6 , 12 and [ \u02c6 M N ] 12 , 12 are presented in Appendix B . The rank of [ \u02c6 M N ] 6 , 12 is rank ( [ \u02c6 M N ] ) = 9 , thus the mobility of the PKM family mechanism is: F N = F \u2212 rank ( [ \u02c6 M N ] ) = 12 \u2212 9 = 3 (28) and the number of redundant constraints is: C N = C N = \u03bb\u03bd \u2212 rank ( [ \u02c6 M N ] ) = 6 \u00b7 2 \u2212 9 = 3 (29) All valid actuation schemes for the PKM family mechanism are presented in Table 3 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001296_aim.2019.8868485-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001296_aim.2019.8868485-Figure7-1.png", + "caption": "Figure 7. Data acquisition method", + "texts": [ + " 2 1 ( ) ) n i i i E Min P Q R T = = \u2212 + (8) In formula (8), P and Q represents two point sets. R and T are the rotation matrix and translation matrix. E is the minimum distance between two point sets. In this paper, ICP algorithm is used to match shapes of objects. According to the matching results, we innovatively apply the ICP to the pose estimation of objects. Besides, in order to obtain the point cloud information of the aero-engine blade using tactile sensor quickly, we propose a rapid measurement method. As shown in Fig.7, the method of fast searching blade edge position is mainly divided into the following four steps. Step one: Control the tactile sensor to detect three random points on the blade. The detected three points can roughly determine the plane of the blade, as shown in Fig.7-(a). Step two: Control the tactile sensor to move a blade length distance along the plane of the blade away from the turbine shaft. Ensuring that the sensor is separated from the area where the blade is, as shown in Fig.7-(b). Step three: Control the sensor to move back along the blade plane until it touches the blade again. And blade edge points can be obtained, as shown in Fig.7-(c). Step four: The point cloud model of the blade can be obtained after sampling along the blade edge, which is shown in Fig.7-(d). The sensor has a hemispherical-shell structure. The outer diameter is 18mm, the thickness is 1mm, and 8 measuring electrodes are uniformly distributed around. The height of the electrodes is 1mm, the width is 1mm, and the thickness is 1mm. Fig.8 shows the whole process of tactile sensor simulation, including physical modeling, grid generation and data measurement. Simulation experiments mainly verify the accuracy of tactile sensor and the effectiveness of the algorithm. In order to obtain the position measurement information of the tactile sensor accurately, the central position of the compression area was first extracted from the imaging image, and then the coordinate information of the compression point was restored", + "14 shows the tactile imaging images obtained by pressing the sensor at different positions. The white cross center in the figure represents the position of the contact center. The errors of different measurement positions are 0.8mm,0.9mm,0.7mm and 0.8mm. The average error relative to the size (diameter of 18mm) of the sensor is 4.4%. B. Blade pose estimation experiment Fig.15 shows the process of data acquisition for Aero-engine blades using AEIR. According to the data acquisition method shown in Fig.7, AEIR is controlled for data acquisition, and the aero-engine blade point cloud model can be obtained. Fig.16 shows the registration graph obtained by ICP algorithm for point cloud information of the blade after data collection. The red dots represent the data points collected, and the blue dots represent the standard model. After matching, the pose information of the blade can be obtained. C. Motion control experiment Fig.17 shows the motion control experiment of the continuum robot. In this experiment, the AEIR is controlled to move along an arc path" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000095_s0025654418050047-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000095_s0025654418050047-Figure7-1.png", + "caption": "Fig. 7.", + "texts": [ + "* = + \u0394 \u03c9 = \u03c9 + \u0394\u03c9 \u03bd = \u03bd + \u0394\u03bd0 0 0, ,a a a \u03c90 0,a \u03bd0 + = ,AZ BZ 0 \u0394 \u03b2 + \u03b4 \u2212 \u03c9 \u2212 \u03c9 \u0394\u03c9 \u03c9 \u03b4\u03c9 + \u03b2 \u03c9 \u2212 \u03bd \u03b2 + \u03b4 \u2212\u03b2 = = = \u0394\u03bd \u03b2 \u03bd \u2212 \u03c9 \u2212\u03b2 \u03b3 + \u03b2 \u0394 2 0 0 0 0 0 0 0 0 0 0 2 2 0 0 0 0 0 0 0 1 1 2 0 0 2 0 0 ( ) ( ) 0 , , . 0 0 1 0 2 ( ) 0 1 0 0 0 0 0 0 1 a a a a a a a a a Z A B MECHANICS OF SOLIDS Vol. 53 Suppl. 2 2018 The coefficients of the characteristic polynomial of the system are (5.4) Analysis of the Routh\u2013Hurwitz conditions shows that they are satisfied for all values of the parameters that satisfy the conditions of existence (5.3). On the left of Fig. 7, the regions of the existence of the auto-oscillatory modes in the plane of the parameters \u03b4 and \u03b2, which follow from conditions (5.3), are shown for different values of the ratio \u03bc/\u03b3. = = \u03b2 + \u03b4 + \u03b3 + \u03b2 = + \u03b2 + \u03b4 + \u03b3 \u03b2 + \u03b4 + \u03b2 \u03b2 + \u03b4 = \u03b3 + \u03b2 + \u03b4 + \u03b2 + \u03b2\u03b4 + \u03b4 = \u03b2\u03b4 0 2 1 0 2 2 2 0 2 2 2 3 0 2 4 0 1, 2( ) , 4 ( ) 2 ( ) ( 2 ) , (4 ( ) ) (4 ) , 4 . k k a k a k a k a = \u2212 > = \u2212 >2 1 1 2 0 3 2 3 1 1 40, 0R k k k k R k R k k MECHANICS OF SOLIDS Vol. 53 Suppl. 2 2018 The areas of stability, which are by shading, are located above the corresponding straight lines. On the right part of Fig. 7 on the AMC graph, the area is shown in a dark color; this portion corresponds to the stable self-oscillation modes. The AMC sections shown by dashes correspond to unstable synchronous whirling. In Fig. 8, the results of the numerical integration of the systems (3.1) and (5.1) are shown, obtained with the following parameter values: The curves 1 correspond to a balanced rotor (\u03b5 = 0), and the curves 2 correspond to an unbalanced rotor (\u03b5 = 0.05). It can be seen from the graphs that the interaction of the self-oscillations and forced oscillations caused by rotor imbalance is expressed in the appearance of high-frequency oscillations of the whirling amplitude, as well as the f luctuations of the whirling speed and the spin speed" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003237_acc45564.2020.9147530-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003237_acc45564.2020.9147530-Figure1-1.png", + "caption": "Figure 1: Depiction of a General Quadrotor", + "texts": [ + " All quaternions representing attitude or rotation are unit quaternions; meaning that they are all subject to the condition: qT q = q20 + qTv qv = 1 (2) The inverse of a quaternion is q\u22121 = q\u2217 qT q (3) where q\u2217 = [q0 -qv]T denotes the conjugate of the quaternion. The time derivative of any quaternion is given by, q\u0307 = 1 2 q\u03c9\u0304 (4) where \u03c9 is the rate of rotation and v\u0304 = [0 vT ]T represents a quaternion equivalent of a vector v \u2208 R3 [2], [6]. Throughout the paper, we denote the rotation matrix corresponding to quaternion q by Rq: Rqv = qv\u0304q\u22121,\u2200v \u2208 R3 (5) Let A = {ea,x ea,y ea,z} represent the aircraft body or local frame of the quadrotor, and I = {ex ey ez} be the inertial or global frame, as shown in Fig. 1. Throughout the paper, we shall also denote the canonical basis vectors, [1 0 0]T , [0 1 0]T , and [0 0 1]T , by ex, ey , and ez , respectively, when no confusion is likely to arise. The dynamic model of the quadrotor, according to the two frames, is given as below [3]: x\u0307 = v (6) mv\u0307 = ftRqez \u2212 gez (7) q\u0307 = 1 2 q\u03c9\u0304b (8) If \u03c9\u0307b = \u03c4a + \u03c9b \u00d7 If\u03c9b \u2212Ga (9) where x \u2208 R3 and v = x\u0307 represent the linear position and velocity vectors, respectively, of the origin of the aircraft frame A with respect to the inertial frame I, q = [q0 qv] T denotes the quaternion representation of the aircraft attitude, \u03c9b = [\u03c9b,1 \u03c9b,2 \u03c9b,3]T represents the angular velocity of the quadrotor relative to the aircraft frame A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001549_s11223-019-00110-8-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001549_s11223-019-00110-8-Figure2-1.png", + "caption": "Fig. 2. Analysis diagram of the relative sliding speed.", + "texts": [ + "1, r1 is the radius of the movable tooth, FHi is the force of the wave generator on the movable tooth, and O1 is the spin angular velocity of the movable tooth. 2.1.1. Kinematic and Force Analysis of Meshing Pairs. When the input wave generator rotates at a constant angular velocity H , the separator has output rotation at a constant angular velocity G . The absolute motion of movable teeth is composed of two motions: along with the uniform rotation of the separator and the relative rotation of the swing-rod around the pin shaft, as shown in Fig. 2. The spin angular velocity of the movable tooth is O A H A O O A r R V 1 1 1 2 1 1 [ cos cos( )], (3) where VO G H G H H G H Ac l k Z Z ckZ 1 2 2 2 2 2 2 2 2 2 sin ( ) cos( )sin( ).ZH The movable tooth and ring gear slide relatively at point B. The speed of B in the tangent direction on the movable tooth is the sliding speed of the movable tooth relative to the ring gear: V B t AO B OV r 1 1 1cos (cos ) sin ( ) cos( B G H G H H G Hc l k Z Z ckZ1 2 2 2 2 2 2 2 2 2 A HZ)sin( ), (4) where A is the included angle of V AO1 and VA , RA is the polar diameter at the point A of the actual contour of the wave generator, O1 is the angle between V O1 and the x-axis, H is the angle between the swing-rod and the x-axis, A is the argument of point A on the actual contour of the wave generator, and ZH is the teeth number of the wave generator, ZH 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003143_jomms.2020.15.291-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003143_jomms.2020.15.291-Figure6-1.png", + "caption": "Figure 6. Deformation processes of self-folding structure from 2D to 3D shapes. Arrow 1: plate compressed at high temperature, Arrow 2: plate cooled down and unloaded. Arrow 3: yellow hinge region reheated to form 3D structure.", + "texts": [ + " In this section, we design a self-folding structure using SMP materials and simulate the deformation processes of the self-folding structure from its two-dimensional plate strip to a three-dimensional structure. The initial configuration of the self-folding structure is a two-dimensional rectangular plate strip shown in Figure 5. The plate thickness comprises two layers of the same shape memory polymer material. The bottom layer in dark blue covers the whole board while the upper SMPs layer is divided into several regions of blue portions connected intermittently by smaller yellow parts of self-folding hinge regions which can later be heated locally. Figure 6 illustrates the deformation processes transforming a two-dimensional pattern to a threedimensional configuration. First, we heat the original plate structure to a high temperature and apply a compressive deformation along the length of the plate. While retaining the deformed shape, the temperature of the whole structure is reduced to ensure that the deformation is frozen to keep the structure in the deformed configuration even after unloading. Finally, while keeping the temperature of the blue and dark blue regions constant, the yellow self-folding hinge regions are heated locally to extend the latter regions to their original shapes. As the deformation of the bottom plate restricts the expansion of the yellow 3D PHASE-EVOLUTION-BASED THERMOMECHANICAL MODEL OF SHAPE MEMORY POLYMER 303 portions, the two-dimensional plate will fold into a three-dimensional pentagon as illustrated in Figure 6. This simple example demonstrates the potential of SMPs for self-folding applications. Many complex structures can be formed or constructed through the rational design and facilitated via the proposed numerical simulation processes. In this paper, we have developed a 3D phase-evolution-based thermomechanical constitutive model for SMPs by extending our previous 1D formulation to enable the analysis of the deformation mechanism at the structural level. The 3D model has been implemented into a finite element package ABAQUS via subroutine UMAT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002696_s40430-020-02368-5-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002696_s40430-020-02368-5-Figure6-1.png", + "caption": "Fig. 6 Finite element model of the gearbox housing", + "texts": [ + " (12)TA = TB = [ Px sin + ( Pz cos )] L 7 (13) A = MzAz0 ( B\u22152 + b 1 ) JzA + MyAy0(H\u22152 + h) JxA \u2212 FnA SA (14) B = MzBz0 ( B\u22152 + b 1 ) JzB + MyBy0(H\u22152 + h) JxB \u2212 FnB SB (15) A = TAh JA (16) B = TBh JB Although the force arm of A\u2013A is shorter than B\u2013B, A\u2013A should be taken as a region of stress concentration. Rough location of stress concentration region could be obtained by force analysis. However, it is difficult through force analysis to get the location and stress distribution of stress concentration region exactly. A finite element model is established for analyzing regions of stress concentration and natural characteristics of gearbox housing (Fig.\u00a06). The model includes 178,123 elements (Solid 185 in ANSYS) and 312,564 nodes. The Young\u2019s elastic modules, yield limit, Poisson\u2019s ratio and density of housing were 200 GPa, 300\u00a0MPa, 0.3 and 7800\u00a0kg/m3. Full constraints were set on the holes of upper and lower hinged ears. A center point was applied for applying loads on the gearbox. External excitation consists of cutting resistance ( Pz ), propulsive resistance ( Px ), axial force ( A 1 ), torque ( Mp ) and gravity (G-including transmission system), as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003210_012011-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003210_012011-Figure2-1.png", + "caption": "Figure 2. Motion referred to generalized body axes.", + "texts": [ + " The airplane main structure is mostly made from composite, gains benefits in lightweight control and enhances the strength of the structure. With retractable landing gears, the airplane can take-off and landing on a short airfield or water runway. A more powerful 115 Hp, Rotax 914 piston engine with three blades constant pitch propeller is installed as \u201cpusher\u201d high-wing configuration. The airplane basic design performance is shown in Table 1. The flying motion of the aircraft analyzing at the center of gravity cg coincident with the origin \ud835\udc5c in an orthogonal axis set (\ud835\udc5c\ud835\udc65\ud835\udc66\ud835\udc67), in the general body, as shown in Figure 2. Then, the force and moment equation comprised the generalized six-degree of freedom equations of motion. The uniform mass distribution of an airframe can be expressed as the following equations: \ud835\udc5a(?\u0307? \u2212 \ud835\udc5f\ud835\udc49 + \ud835\udc5e\ud835\udc4a) = \ud835\udc4b\ud835\udc4e + \ud835\udc4b\ud835\udc54 + \ud835\udc4b\ud835\udc50 + \ud835\udc4b\ud835\udc5d + \ud835\udc4b\ud835\udc51 \ud835\udc5a(?\u0307? \u2212 \ud835\udc5d\ud835\udc4a + \ud835\udc5f\ud835\udc48) = \ud835\udc4c\ud835\udc4e + \ud835\udc4c\ud835\udc54 + \ud835\udc4c\ud835\udc50 + \ud835\udc4c\ud835\udc5d + \ud835\udc4c\ud835\udc51 \ud835\udc5a(?\u0307? \u2212 \ud835\udc5e\ud835\udc48 + \ud835\udc5d\ud835\udc49) = \ud835\udc4d\ud835\udc4e + \ud835\udc4d\ud835\udc54 + \ud835\udc4d\ud835\udc50 + \ud835\udc4d\ud835\udc5d + \ud835\udc4d\ud835\udc51 \ud835\udc3c\ud835\udc65?\u0307? \u2212 (\ud835\udc3c\ud835\udc66 \u2212 \ud835\udc3c\ud835\udc67)\ud835\udc5e\ud835\udc5f \u2212 \ud835\udc3c\ud835\udc65\ud835\udc67(\ud835\udc5d\ud835\udc5e + ?\u0307?) = \ud835\udc3f\ud835\udc4e + \ud835\udc3f\ud835\udc54 + \ud835\udc3f\ud835\udc50 + \ud835\udc3f\ud835\udc5d + \ud835\udc3f\ud835\udc51 \ud835\udc3c\ud835\udc66?\u0307? + (\ud835\udc3c\ud835\udc65 \u2212 \ud835\udc3c\ud835\udc67)\ud835\udc5d\ud835\udc5f + \ud835\udc3c\ud835\udc65\ud835\udc67(\ud835\udc5d 2 \u2212 \ud835\udc5f2) = \ud835\udc40\ud835\udc4e + \ud835\udc40\ud835\udc54 + \ud835\udc40\ud835\udc50 + \ud835\udc40\ud835\udc5d + \ud835\udc40\ud835\udc51 \ud835\udc3c\ud835\udc67?\u0307? \u2212 (\ud835\udc3c\ud835\udc65 \u2212 \ud835\udc3c\ud835\udc66)\ud835\udc5d\ud835\udc5e + \ud835\udc3c\ud835\udc65\ud835\udc67(\ud835\udc5e\ud835\udc5f \u2212 ?\u0307?) = \ud835\udc41\ud835\udc4e + \ud835\udc41\ud835\udc54 + \ud835\udc41\ud835\udc50 + \ud835\udc41\ud835\udc5d + \ud835\udc41\ud835\udc51 (1) Assuming that the disturbing forces and moments on the right side of the equation are the effects resulting from aerodynamic, gravitational, aerodynamic flight control moment, power from engine, and atmospheric disturbances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure8-1.png", + "caption": "Fig. 8. Flux density in the core of stator and rotor of original model", + "texts": [ + "674042 Tesla The distribution of air gap normal flux density at different mechanical rotor positions of the studied PMM can be predicted by numerical analysis based on the finite element, as shown in Fig. 7. It is noted that the air gap magnetic density Authorized licensed use limited to: City, University of London. Downloaded on July 10,2020 at 12:04:30 UTC from IEEE Xplore. Restrictions apply. flux wave is a little bit distorted. It is caused by the slot opening in the core of the stator of all machines. Figure.8 shows that the slot at the edge of the magnet rotor in the Proposed Model and Model 1 (one-step model) does not affect the changing of the variation rate of magnetic flux density wave for all experimental machines. This means that the proposed machine Model is promising for the presence of two steps of slotting in the magnets does not distort the balance of magnetic force in the air gap of the machine of the proposed Model. A PMMs machine under eccentricity, which is the magnetic force in the air gap is unbalanced, and the variation rate of air-gap magnetic flux density is more frequent compared with a health PMMs machine [13]. The magnetic flux density in the core of the stator and rotor investigated using finite element simulation. The maximum values in the core of stator are around 1.4839 Tesla for the Initial Model. For the reason of clear demonstration, three of the maximum magnetic flux density in the stator teeth shown in Fig.8. Simulation results show that the slotting in the magnet edge does not affect the magnetic density flux in the core of the rotor and stator. However, the proposed model has a little bit higher density flux in the stator core compared with model 1. For the magnetic flux density in the rotor core, the proposed Model has the highest around 1.2 Tesla among three machines structures, and Initial Model is 1.4839 Tesla and for Model 1 and proposed Model is 1.20813 Tesla, 1.20968 Tesla, respectively. Considering the limit of density flux magnet in the core of the rotor is around 1.5-1.6 Tesla, it can be noted that the presence of slot at the edge of the magnets may affect to decreasing of magnetic flux density in the rotor core. If the density magnet flux in the core of the rotor above 1.6 Tesla, the performance of the machine is bad and the cross-section of the core must be resized. The density flux in the stator shown in Fig. 8 can be observed that the slot at the edge of the magnet rotor does not affect or only a little bit the increasing of magnetic flux density in stator core and rotor core for the three machines. Based on the fact, it can be concluded that the proposed model can be accepted. It should be noted that the material of stator and rotor core this paper is M-19 which can operate at about 1.5-1.6 Tesla. By using the finite element analysis, the cogging torque was investigated within 180 mechanical degrees with 1(one) degree every step rotor rotation. For clear demonstration, only 60 rotor mechanical degrees are depicted in Fig 8. It was found that the cogging torque for all machines has two pulsations for every 600 rotors mechanical degree. The CT for Initial Model is around 0.02 N-m (peak value). It is acceptable since the original model has the largest air-gap reluctance among the tested models. The CT peak for model 1 is 0.015 N-m, which is lower compared with Initial Structure. The decrement sourced from the effect of rotor slotting in magnet and the air-gap reluctance. The CT peak for the proposed model is 0.01 N-m; it is the lowest torque among three experimental models" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003807_icuas48674.2020.9214047-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003807_icuas48674.2020.9214047-Figure1-1.png", + "caption": "Figure 1. Jet engine powered UAS (quad-jet) free body diagram", + "texts": [ + " The experimental studies and the developed prototype have been presented in Section V. Concluding remarks and future work are presented in Section VI. Before explaining the mathematical model of the quad-jet system, the coordinate systems must be defined. During the modeling, 6 coordinate systems will be used. The first is the inertial reference system, whose unit vectors are oriented to north-east-down and named as INED. The second reference system is located at the center of mass of the vehicle. The center of the moving reference system shown in Figure 1 is defined as B. This reference system allows the forces and moments to be acted on by the engines to be easily formulated. The other four reference systems are also moving reference systems, placed in positions where the center of mass of the jet engines are located and aligned with the vectors of the B axis. Euler angles are used to mathematically express the angular orientation of the moving reference systems used in the aircraft. In this study, ZYX Euler angles representation was preferred. The symbol \u03c8 is used for rotation around the Z axis, \u03b8 for rotation around the Y axis, and \u03d5 for the rotation around the X axis", + " Considering Newton's law of motion and Euler equations, the relationship between forces and moments affecting the aircraft and the motion of the vehicle is obtained with the help of equations (3) and (4). { \ud835\udc39\ud835\udc35\ud835\udc65 \ud835\udc39\ud835\udc35\ud835\udc66 \ud835\udc39\ud835\udc35\ud835\udc67 } = \ud835\udc5a ({ \ud835\udc5d \ud835\udc5e \ud835\udc5f } \u00d7 { \ud835\udc62 \ud835\udc63 \ud835\udc64 } + { ?\u0307? ?\u0307? ?\u0307? }) { \ud835\udc40\ud835\udc35\ud835\udc65 \ud835\udc40\ud835\udc35\ud835\udc66 \ud835\udc40\ud835\udc35\ud835\udc67 } = [ \ud835\udc3c\ud835\udc65 0 0 0 \ud835\udc3c\ud835\udc66 0 0 0 \ud835\udc3c\ud835\udc67 ] . { ?\u0307? ?\u0307? ?\u0307? } + { \ud835\udc5d \ud835\udc5e \ud835\udc5f } ([ \ud835\udc3c\ud835\udc65 0 0 0 \ud835\udc3c\ud835\udc66 0 0 0 \ud835\udc3c\ud835\udc67 ] . { \ud835\udc5d \ud835\udc5e \ud835\udc5f }) Here FB and MB represents forces and moments acting on the vehicle and represented in B reference frame. The aircraft is assumed to be symmetric with respect to the B reference frame, resulting a diagonal inertia matrix. The free body diagram of the quad-jet is displayed in Figure 1. Total force and the moment vector acting on the quad-jet will be found. Each motor can be rotated around y and x axes as shown in the Figure 2, resulting a change in the direction of the thrust of the engine. The rotation in each motor is represented by Euler angles i and i i.e. (i: 1, 2, 3, 4). Here, again ZYX convention is used, and the rotation around Z is taken as zero. 876 Authorized licensed use limited to: Newcastle University. Downloaded on October 18,2020 at 15:29:54 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003429_0954407020948397-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003429_0954407020948397-Figure6-1.png", + "caption": "Figure 6. A car-like vehicle and its kinematic model.", + "texts": [ + " Geometric constraints are enforced only on the position o of the system fi(o)=0, i=0, 1, . . . ,m\\ n \u00f04\u00de Kinematic constraints are represented as analytical relations between the position o and the velocity _o of the system gi(o, _o)=0, i=0, 1, . . . ,m\\ n \u00f05\u00de Kinematic constraints (equation (5)) that cannot be integrated to the form of equation (4) are called nonholonomic constraints. Systems subject to nonholonomic constraints are called non-holonomic systems. Formally, the kinematic constraints to which a carlike vehicle (shown in Figure 6) is subject can be written as74 _x _y _f 2 4 3 5= v sinf v cosf (v tan u)=L 2 4 3 5 \u00f06\u00de where \u00bdxyf T is the pose of the vehicle, v is the longitudinal velocity of the vehicle, u is the steering angle, and L is the wheelbase. A car-like vehicle is an archetypal non-holonomic system. It cannot move sideways, and its turning radius is lower bounded. Equation (6) can be written as a discretized state-transition equation xi yi fi 2 64 3 75 = xi 1 yi 1 fi 1 2 64 3 75+ vi 1 dt sinfi 1 vi 1 dt cosfi 1 (vi 1 dt tan ui 1)=L 2 64 3 75 , si = fT(si 1,mi 1) \u00f07\u00de where mi 1 = \u00bdvi 1ui 1 T is the control vector and dt is the time step" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003696_0954409720962245-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003696_0954409720962245-Figure6-1.png", + "caption": "Figure 6. Vehicle dynamics model.", + "texts": [ + " Vehicle dynamics model featuring asymmetric OOR coupled with wear Dynamics model of vehicle system A vehicle dynamics model was created based on the parameters of a subway train in China. The entire vehicle model consists of 1 vehicle body, 2 frames, 4 wheelsets, 8 axle boxes, and primary and secondary suspension systems. The vehicle body, each frame and each wheelset have 6 degrees of freedom (DOFs), including swaying, surging, heaving, pitching, rolling, and yawing. Each axle box has 1 DOF, i.e. pitching. Then the whole vehicle has a total of 50 DOFs. Table 1 provides the main vehicle parameters, the dynamics model of the vehicle system is shown in Figure 6. The 1200m line was composed of straight lines, transition curves and circular curves. The detailed length of each line type is shown in Table 2, the circular curves were modeled as an S-curve with a radius of R600, as shown in Figure 7. A subway train needs to accelerate, coast, and then decelerate when it passes across a subway station.25 The speeds for subway trains of types B and C are specified in the Construction Standard for Urban Rapid Rail Transit Projects. For consistency with the speed in actual operation, the speed variation induced by traction and braking of the test vehicle was adopted as the basis for setting the speed during simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002215_s00170-019-04806-8-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002215_s00170-019-04806-8-Figure7-1.png", + "caption": "Fig. 7 Projections of self-rotation velocities to moving axes\u041e2\u03b5 and\u041e2\u03b7", + "texts": [ + " V\u03c8 B1\u03b5 \u00bc V\u03c8 B1e \u00bc \u2212S1\u03c8\u02d9 ; V\u03c8 B2\u03b5 \u00bc V\u03c8 B2e \u00bc \u2212S1\u03c8\u02d9 ; V\u03c8 K\u03b5 \u00bc 0:5D\u03c8\u02d9 ; V\u03c8 B1\u03b7 \u00bc V\u03c8 B1hcos\u03b3 \u00bc b\u03c8\u02d9 cos\u03b3; V\u03c8 B2\u03b7 \u00bc V\u03c8 B2hcos\u03b3 \u00bc \u2212b\u03c8\u02d9 cos\u03b3; V\u03c8 K\u03b7 \u00bc 0; V\u03c8 B1\u03b6 \u00bc \u2212V\u03c8 B1hsin\u03b3 \u00bc \u2212b\u03c8\u02d9 sin\u03b3; V\u03c8 B2\u03b6 \u00bc V\u03c8 B2hsin\u03b3 \u00bc b\u03c8\u02d9 sin\u03b3: V\u03c8 K\u03b6 \u00bc 0: \u00f015\u00de Rotation of peg about its axis occurs with angular velocity \u03c6\u0307 \u00bc d\u03c6 dt . Velocities V\u03c6 K , V \u03c6 B1, V \u03c6 B2 of contact points \u041a, \u04121 and \u04122 while moving are located in the plane of aligned peg end \u041e2\u03b5\u03b7, and are equal in magnitude: V\u03c6 K \u00bc V\u03c6 B1 \u00bc V\u03c6 B2 \u00bc 0:5d\u03c6\u02d9 : Projections of these velocities to peg \u041e2\u03b6 axis are equal to zero: V\u03c6 B1\u03b6 \u00bc V\u03c6 B2\u03b6 \u00bc V\u03c6 K\u03b6 \u00bc 0: Projections to moving axes \u041e2\u03b5 and \u041e2\u03b7 (see Fig.7) are equal to. V\u03c6 \u04121\u03b5 \u00bc V\u03c6 B1cos\u03b2 \u00bc \u22120:5d\u03c6\u02d9 S2 0:5d \u00bc \u2212S2\u03c6\u02d9 ; V\u03c6 \u04121\u03b7 \u00bc V\u03c6 \u04121sin\u03b2 \u00bc 0:5d\u03c6\u02d9 b 0:5d \u00bc b\u03c6\u02d9 : V\u03c6 \u04122\u03b5 \u00bc \u2212V\u03c6 B1cos\u03b2 \u00bc \u22120:5d\u03c6\u02d9 S2 0:5d \u00bc \u2212S2\u03c6\u02d9 ; V\u03c6 \u04122\u03b7 \u00bc V\u03c6 \u04121sin\u03b2 \u00bc \u22120:5d\u03c6\u02d9 b 0:5d \u00bc \u2212b\u03c6\u02d9 : V\u03c6 K\u03b5 \u00bc 0:5d\u03c6\u02d9 ; V\u03c6 K\u03b7 \u00bc 0: \u00f016\u00de Each of V\u03c6 B1\u03b7 and V\u03c6 B2\u03b7 shall be divided into two compo- nents, one of which is directed along axis \u041e1h, the other is parallel to axis \u041e1z (see Fig. 8) V \u03c6 B2\u03b7 \u00bc V \u03c6 B2h \u00fe V \u03c6 B2z: Hence, velocity in each point may be presented as a sum of three components, two of which are located in horizontal plane \u041e1\u0445\u0443, and the third one is located vertically, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001822_asyu48272.2019.8946340-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001822_asyu48272.2019.8946340-Figure1-1.png", + "caption": "Fig. 1. Cross section of flip-chip C4 package with CMOS chips.", + "texts": [ + " In such a case, the solution of (14) can yield two results for the quadratic equation, and the larger one is omitted due to its metastable state which would be alternated by small perturbation [28]. Finally, the steady-state temperature can be obtained Tsteady_state = \u2212B \u2212 \u221a B2 \u2212 4AX 2A . (15) It is noted that the value of B in (15) is negative and its absolute value is large than the square root in the real calculation process. There is no doubt that the value of A in (15) is positive, thus the final steady-state temperature Tsteady_state is also positive. The traditional flip-chip C4 package structure is shown in Fig. 1 [3]. The C4 solder bump distributed over the next layer substrate is soldered and connected to the chip die. While, epoxy commonly used on large high-power chips is underfilled between the die and the substrate. Thermal paste, placed between the aluminum lid and the silicon die to establish good thermal contact, is widely used as thermal interface material and can provide protection to the chips from the mechanical damage [10]. The thermal on the powered die chip is spread quickly to the surrounding Al cap via thermal paste" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure87.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure87.2-1.png", + "caption": "Fig. 87.2 3D solid model and FE Meshed model", + "texts": [ + " As the results of thermal analysis was further used to determine the residual stress and strain induced in the plate, thus 87 Thermomechanical Analyses of Single Sided Single Pass \u2026 1035 the eight- node SOLID70 element was later replaced to eight-node SOLID185 element to perform structural analysis. The weld beads obtained from experiments were analyzed using Nikon SMZ25 microscope with magnification of 0.5 and 30 W scanning power. The bead geometry thus obtained was used for developing FE model. The modelling details single sided SAW is shown in Fig. 87.2. The solid model was built based on actual experimental bead cross section, further the model was meshed using ANSYS 18.0. Fine mesh was carried out along the weld region to obtain better temperature gradient along weld zone and to reduce computation time coarse mesh was used away from the bead zone. The 3D solid model and FE mesh model was shown in Fig. 87.2. The temperature dependent material properties [11] of austenitic steel is mentioned in Figs. 87.3 and 87.4. 1036 P. V. S. S. Sridhar et al. The modelling of a transient thermal analysis with arc moving from one location to another with respect to time in order to generate uniform weld bead, the element birth and death technique was used. The entire weld pool was divided into multiple elemental volumes and subsequently with respect to time, heat was added and removed consecutively from one bead to another, thus attaining a moving heat source model" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001711_icems.2019.8922442-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001711_icems.2019.8922442-Figure1-1.png", + "caption": "Fig. 1. Prototype of 6\u2013phase 36S/34P FSCW CP PMSM used for experimental investigations. (a) Test setup. (b) Cross\u2013section.", + "texts": [ + " ( ) ( ) ' ; ;J f J SPC = (18) The learning rate for every parameter \u03b8i at every time step t can be defined as in (19). ( ) ( )1, , , 2 , 1 t i t i t i i t i t J J + = = \u2212 + (19) Upon implementing the adaptive gradient algorithm, the best slot\u2013pole combinations with minimal inductance and induced EMF harmonics for 3\u2013, 5\u2013 and 6\u2013phase machines are obtained and are shown in Table I\u2013III respectively. The UMF for these SPC with minimal inductance and EMF harmonics are identified using algorithm 1. Experimental investigations on a prototype of 6\u2013phase, 36\u2013 slot, 34\u2013pole FSCW CP PMSM as in Fig.1(a) were performed in order to verify the analytical model for inductance and induced EMF. Further, an electromagnetic model for the prototyped machine was analyzed using FEA. The cross\u2013section and detailed parameters of 36S/34P FSCW CP PMSM are shown in Figs. 1(b) and 2(b) respectively. Results obtained from experimental analysis and simulations were compared to the analytical model. A. Verification of Proposed Inductance Model In order to measure inductance for every rotor position, LCZ meter was used for the laboratory CP PMSM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001371_012040-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001371_012040-Figure2-1.png", + "caption": "Figure 2. 3-D finite element model. (a) Wheel/rail-foundation contact system, (b) Detail of contact and (c) Contact pair.", + "texts": [ + " The Solid226, Conta174 and Targe170 have structural-thermal capability. In order to reduce the element and node numbers, the multiple point constraint (MPC) method is used. The model has 48672 nodes and 42101 elements. The smallest element size is 0.5 mm in this article [7,8]. At the same time, the coupling of both heat and force is strong coupling, and the full integration method is adopted. At last, the Solid186 is selected to establish temperature-independent model, in which the temperature is not considered. And calculation model is shown in figure 2. When the rail length exceeds 4.2 m, the rail length has little effect on the result. Therefore, the length of rail in FEM is 4.2 m [8]. The longitudinal and transverse displacement on the two sides of rail are constrained. The vertical displacement at the bottom of foundation springs is constrained. For the fastener springs, the vertical displacement at the top is constrained. The wheel can slide on the rail surface along longitudinal direction, but it can not move in the transverse direction. The wheel load is 100 kN" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003496_ecce44975.2020.9236284-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003496_ecce44975.2020.9236284-Figure1-1.png", + "caption": "Fig. 1. (a) low-cost single-phase BLDC claw-pole motor with a bifilar winding and (b) circuit to drive the bifilar winding [6]\u2013[8].", + "texts": [ + " For verification of thermal analyses, measurements are mostly carried out using temperature sensors (e.g., thermocouple, PT1000) located in different parts of the machine. Automotive applications generally have high temperature requirements. Auxiliary drives must function properly over a wide temperature range, e.g., between \u221240\u25e6C and 135\u25e6C. Hence, knowledge of the thermal performance of the used motors is vitally important. This paper investigates the thermal characteristics of different design variations of the BLDC claw-pole motor in Fig. 1(a) through 3-D finite-element simulations and experiments. To this aim, the steadystate and transient overtemperatures of the winding are analyzed for different input losses dissipated in the 978-1-7281-5826-6/20/$31.00 \u00a92020 IEEE 4269 Authorized licensed use limited to: Carleton University. Downloaded on June 01,2021 at 06:08:18 UTC from IEEE Xplore. Restrictions apply. ring winding of the claw-pole stators. The results are used to determine the parameters of a simple thermal equivalent circuit. The reminder of this paper is structured as follows: Section II reviews the structure and working principle of the low-cost single-phase BLDC claw-pole motor. Section III discusses the techniques used for the thermal analyses. Section IV presents the results of the thermal investigations, and Section V concludes the findings of this paper. II. BLDC CLAW-POLE MOTOR STRUCTURE AND WORKING PRINCIPLE As illustrated in Fig. 1(a), the low-cost BLDC clawpole motor topology consists of an outer-rotor and a bifilar ring winding on a bobbin [which is not shown in Fig. 1(a)] housed by two punched and subsequently deep-drawn steel sheet parts which form the stator [6]\u2013 [8]. The dashed line denoted \u03a6 illustrates the threedimensional flux path in this type of electric motor. The inverter circuit to drive the motor\u2019s bifilar winding is depicted in Fig. 1(b). A Hall sensor is typically used for the commutation of the switches SA and SB, while an asymmetric air-gap enables the starting of this single-phase PM motor. The air-gap radius of the example case motor is about 8 mm, the axial length is approximately 6 mm, the output power is about 0.7 W, and the rotational speed is 5500 rpm. The application is an automotive cooling fan. III. THERMAL ANALYSES This section discusses the techniques used for the thermal analyses, including the used numerical models as well as the experimental test setups", + ", 30\u25e6C and 60\u25e6C, to investigate any possible temperature dependence of the temperature coefficients. 4270 Authorized licensed use limited to: Carleton University. Downloaded on June 01,2021 at 06:08:18 UTC from IEEE Xplore. Restrictions apply. Stator A Stator B Stator C Stator 0 Fig. 3. Finite-element models of the claw-pole stator designs under investigation. Fig. 3 shows the finite-element models of the different claw-pole stators under investigation. Stator A is the baseline claw-pole motor design shown in Fig. 1(a). In Stator B (which has auxiliary slots) the cogging torque is modulated and Stator C (which has auxiliary slots and skewed stator claws) is optimized for low cogging torque [7]. Due to the removal of iron material, the cooling surfaces of all three stators differ. As a worst-case scenario, in the model denoted Stator 0, the claw-poles are taken off completely, see Fig. 3. Fig. 4 shows the thermal finite-element model of the baseline claw-pole stator, i.e., Stator A, implemented in JMAG [12], illustrating the motor components, the used mesh, and a typical thermal contour plot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002002_robio49542.2019.8961771-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002002_robio49542.2019.8961771-Figure1-1.png", + "caption": "Figure 1. Overall sketch of the continuum robot (include three posture sensors and one permanent magnet). The continuum robot consists of 22 cylindrical connecting disks made of resin material and 21 pairs of 5-mm-long NiTi columns. Each connecting disk has a thickness of 5 mm and a diameter of 30 mm. The length of the single section is 215mm.", + "texts": [ + " So what this paper focuses is how to arrange sensors according to the simulation analysis of continuum robot affected by external load to ensure the simplicity of configuration and the accuracy of shape sensing. In order to solve the problem that traditional continuum robot have twisting problems in their central located flexible back-bone, [20] proposes a new structure. In this paper, the posture sensors and magnet are integrated into the continuum robot with this new structure to detect its shape affected by external load. As shown in Figure 1, the continuum robot consists of 22 cylindrical connecting disks made of resin material and 21 pairs of 5-mm-long NiTi columns. Each connecting disk has a thickness of 5 mm and a diameter of 30 mm. The length of the single section is 215mm, which can realize flexible movement in 3D space. Posture sensors mainly include gyroscope sensor, three-axis acceleration sensor and three-axis magnetic sensors. A permanent magnet is installed as a magnetic source at the starting point of a continuous robot" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001941_s12647-019-00363-3-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001941_s12647-019-00363-3-Figure3-1.png", + "caption": "Fig. 3 Transformation of the CCPG physical structure to a 2D finite element model including an example of a convergent solution", + "texts": [ + " The equations to calculate effective area (Ap), distortion coefficient (k) and piston fall rate (m) are given as follows [4]; AP \u00bc pr20 1\u00fe h0 r0 1 r0p Z l 0 uz \u00fe Uz\u00f0 \u00de dpz dz dz 8 < : 9 = ; \u00f02\u00de k \u00bc Ap pr20 1\u00fe h0 r0 1 0 @ 1 A 1 p \u00f03\u00de m \u00bc p r0q\u00f0p\u00de 6 Rl 0 g\u00f0pz\u00de q\u00f0pz\u00de 1 h3z dz \u00f04\u00de where h0 and r0 are the undistorted gap and piston radii, respectively. Equation (3) is suitable only for an \u2018\u2018ideal gap profile\u2019\u2019 used in this study. CCPG was assumed to have an \u2018\u2018ideal gap profile\u2019\u2019 which is perfectly straight and cylindrical. Therefore, a 2D axis-symmetry with a static structural FEA analysis can be modeled using the ANSYS parametric design language (APDL), as shown in Fig. 3. The model used a solid- rectangular with eight nodes of the PLANE 183 element type. It consists of around 7000 meshes, with the different densities concentrated along the P/C engagement which should be investigated. A total of 407 nodes were created along each surface of the P/C. Table 1 shows the dimensional and material properties of the model used in this investigation, as in [4, 8] since the CCPGs are assumed to be identical. The evaluation of uncertainty concerns possible error sources, which can come from the input parameters, such as the dimensional and material properties and boundary conditions of the model [4, 5]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002215_s00170-019-04806-8-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002215_s00170-019-04806-8-Figure1-1.png", + "caption": "Fig. 1 Position of moving coordinate system \u041e2\u03b5\u03b7\u03b6 relative to fixed system \u041e1\u0445\u0443z at the beginning of alignment", + "texts": [ + " It allows to consider the peg-on-hole motion in the most general case, when the forces, transfer all three types of peg-on-hole movements allowed by connections: align, slip, and slide. The main part of the work is devoted to the step-by-step derivation of DDE in Section 3. This includes the study of directions of velocities, normal reactions, and friction forces at the contact points. The goal pursued with DDE in this paper is addressing three major problems of the process for automated assembly of cylindrical parts: 1. Determining patterns of peg movement that facilitates alignment of the parts when all three generalized coordinates (Fig. 1) change: nutation angle \u03b3 = \u03b3(t), precession angle \u03c8=\u03c8(t), and self-rotation angle \u03c6 =\u03c6(t). 2. Determining forces acting on peg at the points of contact and potentially preventing alignment of the parts. Two example cases of solution of these problems were considered: (i) reaction forces for relative large alignment angle during uniform alignment and slip \u03b3\u0307 \u00bc const; \u03c6\u0307 \u00bc const; \u03c8\u0307 \u00bc 0 in comparison to the case without slip \u03b3\u0307 \u00bc const; \u03c6\u0307 \u00bc 0; \u03c8\u0307 \u00bc 0; (ii) the kinematics and directional cosines were studied for the small alignment angle approximation, and the reaction forces were analysed", + " All necessary characteristics of such motion might be obtained applying differential equations of the mass center movement (1) and Lagrange differential Eq. (2). m d2\u0445c dt2 \u00bc \u2211Fkx ; m d2yc dt2 \u00bc \u2211Fky; m d2zc dt2 \u00bc \u2211Fkz; \u00f01\u00de d dt \u2202T \u2202\u03b3\u0307 \u2212 \u2202T \u2202\u03b3 \u00bc Q\u03b3 ; d dt \u2202T \u2202\u03c8\u0307 \u2212 \u2202T \u2202\u03c8 \u00bc Q\u03c8; d dt \u2202T \u2202\u03d5\u0307 \u2212 \u2202T \u2202\u03d5 \u00bc Q\u03d5; \u00f02\u00de where Fkx, Fky, and Fkz are the projections of the external forces k on the axes x, y, and z correspondingly. \u0422 is a kinematic energy of the peg during the motion,Q\u03b3,Q\u03c8, andQ\u03c6 are generalized forces of Lagrange. Three coordinate systems are canonical to use for these three motions of the system: (i) fixed system of coordinates \u041e1xyz (see, e.g., Fig. 1), which are used in the definition of Lagrange equations; (ii) coordinate system \u041e1hez (see, e.g., Fig. 2) performing precession motion \u03c8 with the peg around hole axis, in which peg could be considered without precession; and (iii) moving coordinate system \u041e2\u03b5\u03b7\u03b6 associated with peg, performing nutation motion \u03b3. In this system, the self-rotational motion of peg around \u041e2\u03b6 is considered. These coordinate systems are used to consider three generalized motions independently. Mathematically, transitions from (iii) to (ii), and from (ii) to (i) include nutation and precession motions into consideration consequently", + " Thus, we use coordinates h and z to describe the dynamics and geometry in DDE (29). Now, the left and right parts of (1) and (2) will be considered separately in order to derive DDE (29). The left parts of the first three differential equations (1) are the time derivatives from coordinates \u0445c, \u0443c, and zc of the peg mass center that are the functions of generalized coordinates \u03b3, \u03c8, and \u03c6. The peg position in the process of aligning is defined relative to the fixed system of coordinates \u041e1xyz (see Fig. 1), the beginning of which coincides with center \u041e1 of the hole aperture edge, axis \u041e1z is directed along the hole axis, coordinate plane \u041e1xz passes through the hole axis, and the peg axis in the initial position of the parts, axis \u041e1x is the line of intersection of this plane with the horizontal plane of the edge, it coincides with the hole diameter, axis \u041e1y is perpendicular to plane \u041e1xz. A moving coordinate system \u041e2\u03b5\u03b7\u03b6 (Fig. 1) is associated to the peg, with the beginning at the center of its aligned end, axis \u041e2\u03b6 is directed along the peg axis, coordinate plane \u041e2\u03b6\u03b7 passes through axes of the parts and is a principal plane of their symmetry in the course of alignment, axis \u041e2\u03b7 coincides with the diameter of a peg located in the plane of symmetry, axis \u041e2\u03b5 is perpendicular to plane \u041e2\u03b6\u03b7. When the peg moves, axis \u041e2\u03b5 remains parallel to segment \u04121\u04122, axis \u041e2\u03b7 remains perpendicular to this segment. To define a position of peg during its rotation about the sleeve axis, an auxiliary system of coordinates \u041e1\u0435hz is used, with the beginning at point \u041e1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001297_s11668-019-00756-1-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001297_s11668-019-00756-1-Figure6-1.png", + "caption": "Fig. 6 Dynamic model of rolling bearing with absence of one rolling body", + "texts": [ + " The theoretical results were basically consistent with the simulation results, and the effectiveness of simulation model was validated by the small relative error of 4.8%. Fracture Fault of Rolling Body In this part, the case of rolling body fracture during operation is selected as analysis object. To simplify analysis, fracture fault of rolling body is handled with extreme form; namely, rolling body is missing, and dynamic model of a deep groove ball bearing with lack of rolling body is established, as shown in Fig. 6. Figure 7 shows effective stress simulation results of bearing rolling bodies are extracted at 0.02 s based on the model built in Fig. 7. Similarly, effective stress results of rolling with the fracture of two rolling bodies at 0.02 s are shown in Fig. 8. As is shown by Figs. 7 and 8, contact stress distribution of rolling bodies with the absence of one or two rolling bodies is the same as that of rolling bodies in fault-free bearing under normal circumstances, but compared to the maximum stress of 2" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure77.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure77.2-1.png", + "caption": "Fig. 77.2 Schematic of spur gear", + "texts": [ + " The input to the motor and the field of the alternator is controlled externally by two separate autotransformers. The autotransformer is used to adjust speed and torque. There is an additional rectifier circuit consisting of diodes and capacitor for the AC to DC conversion. A 100 X rheostat is used as the load across the alternator. The testing gear is connected to the alternator shaft while the standard steel gear is connected to the motor shaft. Evaluation of Adhesive Spur Gear Performance. Figure 77.2 shows the schematic of spur gear used for fabrication. Table 77.1 shows the parameters of the adhesive spur gear (Anand Mohan and Senthilvelan [12]). The adhesive mixture with different properties by varying hardener to resin (H/R) ratio like 0.6:1, 0.7:1, 0.8:1, 0.9:1 and 1:1 was cast into the suitable mould with a thickness of 6 mm (Fig. 77.3a). The cast adhesive mixture into the mould was allowed to cure for more than 24 h and removed from the mould carefully. Figure 77.3b shows fabricated spur gear made of epoxy adhesive system and a steel gear for a reference" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000303_978-981-13-3627-0-Figure5.23-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000303_978-981-13-3627-0-Figure5.23-1.png", + "caption": "Fig. 5.23 Finite element model for the impact effect analysis of the needle ejecting process: a without penetration of the substrate, b with through-penetration of the substrate. \u00a9 2011 IEEE. Reprinted, with permission, from Ref. [39]", + "texts": [ + " The geometry dimensions, loading conditions, and material properties of the finite element model are listed in Tables 5.2 and 5.3, which represent a common case of the chip-on-substrate structure for the peeling-off process in electronics assembly. In consideration of the industrial practice, there may exist two different situations for the analysis of the impact effect by the needle ejecting, that is, without and with penetration of the substrate. The finite elementmodel for these two cases are depicted 132 5 Single-needle Peeling in Fig. 5.23. Local regions surrounding the contact points are refined to ensure the accuracy of numerical results. The dynamic impact force acts directly on the chip in the through-penetration case, resulting in a high instantaneous stress concentration at the local contact area on the backside of the chip. The impact speed is an important factor affecting the stress level. The VonMises stress contour of the chip subjected to the impact force induced by the ejector needle at a speed of 0.1 m/s under the through-penetration condition is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001555_smc.2019.8913935-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001555_smc.2019.8913935-Figure1-1.png", + "caption": "Fig. 1. The coordinate and relative position vector", + "texts": [ + "e, \u2225\u2225ri(t)\u2212 r j(t) \u2225\u2225< 200m, and the initial positions are given by r0 = [450,134,260]m, r1 = [310,110,195]m, r2 = [305,98,340]m, r3 = [155,46,96]m, r4 = [167,53,410]m. The UAV\u2019s dynamics is defined in [30] and the leader\u2019s trajectory is as follows V0 = 80+8sin(0.16t) \u03b80 = 0.12sin(0.04t) \u03c80 = 0.02sin(0.1t) (15) where V0 is the speed, \u03b80 is the flight-path angle, \u03c80 is the heading angle. To maintain the formation shape, each UAV needs to maintain a relative position vector with the leader, as shown in Fig. 1. Then, we can obtain that ri(t)\u2212hi\u2192 r0(t) in finite time. The control protocol (7) will be rewritten as ui (t) =\u2212\u03b2 N \u2211 j=1 ( ai j +\u03c8 (\u2225\u2225epi\u2212 ep j\u2212hi j \u2225\u2225))sig(epi\u2212 ep j\u2212hi j) \u2212\u03b2 ( ai0 +\u03c8 (\u2225\u2225epi\u2212hi \u2225\u2225))sig(epi\u2212hi) \u03b11 \u2212 \u03b3 N \u2211 j=1 ai jsig(eqi\u2212 eq j) \u03b12 \u2212 \u03b3ai0sig(eqi) \u03b12 + \u03be\u0307i (t) (16) The desired formation vectors are given by h1 =[ 100 50 \u2212100 ]T , h2 = [ 100 50 100 ]T , h3 =[ \u2212220 \u221250 \u2212100 ]T , h4 = [ \u2212220 \u221250 100 ]T . The observer parameters are chosen as k1 = 5, k2 = 4.5, c1 = 9, d1 = 7, c2 = 7, d2 = 9" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000682_ijsi-10-2018-0076-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000682_ijsi-10-2018-0076-Figure7-1.png", + "caption": "Figure 7. (a) Input boundary conditions and (b) meshing details", + "texts": [ + " As we approach the highest contact point between the two contact surfaces (Y -cord \u00bc 10\u22124), a significant deflection in normal mechanical stress between the two cases is observed. In the frictionless case (sliding surfaces), thermal load is neglected; therefore, stress is fully translated into mechanical (shear), whereas in the case of \u03bc \u00bc 1 (sticking), a great amount of energy is converted into thermal stresses (which are not presented in the current work), justifying the lower mechanical stress value and denoting the pure contact in dead centers of an internal combustion engine. Figure 7 shows the simulation model for ring\u2012liner contact. The first step of every finite element model is to validate an existing theoretical model. Therefore, an analysis should be conducted in the absence of any axial forces, considering only the radial ones acting on the ring. The number of elements used is 38.410 and the number of nodes is 38.577. The convergence is achieved using CONTACT172 and TARGE169 type of elements. The number of nodes that are in sliding denotes the contact surfaces. In Figure 7(b), the grid of nodes for one of the solutions obtained is indicated. It is obvious that the nodes that are in sliding condition, and therefore in pure contact, define, thus, the contact surface. Dividing the stresses in x-direction over the stresses in y-direction, the kinematic friction coefficient in the certain stresses angle can be calculated. These values are used later on as inputs in the contact model. 5\u00d7107 4\u00d7107 3\u00d7107 2\u00d7107 1\u00d7107 0 0.0 0.2 0.4 0.6 0.8 1.0 Normalised piston ring width (m) FE normal stress (friction coef=1) Analytical Herzian solution (Hertz, 1896) FE normal stress (friction coef=0) N or m al st re ss v er tic al to c on ta ct su rf ac e (P a) Figure 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001699_1350650119893896-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001699_1350650119893896-Figure8-1.png", + "caption": "Figure 8. Schematic diagram of the nonlinear stability boundary.", + "texts": [ + " In the present study, the same technique is introduced to find out the nonlinear stability boundary of the finite-length hybrid journal bearing. Basically, the larger the distance between the initial point and equilibrium point is, the greater the likelihood of the unstable state will be. There must be a critical point in a certain direction that separates the stable part and the instable one. All the critical points connected with each other can form the nonlinear stability boundary. Any initial points located inside the boundary lead to convergent orbits, while the points outside the boundary make the system unstable. Figure 8 is the schematic diagram of the nonlinear stability boundary. Os is the stable equilibrium point, and Ob is the bearing center. The clearance circle with the radius of one is the bearing clearance. The clearance circle boundary is the limited range of the motion of the shaft center, and the points on the clearance circle are unstable points. A critical point should exist on each line from the equilibrium point Os to the corresponding point on the clearance circle. The Ocritical, for example, is the critical point on the line OccOs in Figure 8. The circle boundary is divided into multiple parts with uniform angles. In the present study, the circle is decomposed into 24 parts. Through connecting the critical points on each part, the nonlinear stability boundary can be obtained. Figure 9 shows the method of searching the critical point in each part. The details are demonstrated as follows: 1. The Occ is assumed as one of the points on the clearance circle. Firstly, pick the middle point O1 of the line OSOcc. 2. Calculate the orbit of shaft center with the initial point located at O1 and using equation (27) to determine the orbit state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001583_s00202-019-00889-4-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001583_s00202-019-00889-4-Figure8-1.png", + "caption": "Fig. 8 Machine topology", + "texts": [ + " An incremental encoder provides the rotor position and mechanical speed to determine the reference angular position sector and estimate the back-emf. With this estimation and measuring the line current, the instantaneous P-Q power can be obtained and compared with the power reference values. The reference of active power is determined by a PI controller (control of VDC ) and the reference of reactive power is set to zero for unity power factor operation. Finally, a commutation vector is generated to obtain the needed power variation in each execution cycle of the algorithm. Figure\u00a08 shows the negative-saliency AFPMSG used in this work. The machine has a double rotor and central stator configuration [2]. (9) P\u2217 \u2212 \ud835\udf00P 2 > Pgen \u2192 P + + P\u2217 + \ud835\udf00P 2 < Pgen \u2192 P \u2212 \u2212 ( P\u2217 + \ud835\udf00P 2 \u2265 Pgen P\u2217 \u2212 \ud835\udf00P 2 \u2264 Pgen ) \u2192 P == Q\u2217 \u2212 \ud835\udf00Q 2 > Qgen \u2192 Q + + Q\u2217 + \ud835\udf00Q 2 < Qgen \u2192 Q \u2212 \u2212 The particular negative-saliency characteristic is achieved by replacing with soft iron a portion of the magnet material in the rotor poles (Fig.\u00a09a). Therefore, the d-axis inductance is increased, while q-axis inductance is almost not affected, leading to the condition that Ld is higher than Lq (negative saliency) corresponding to the inverse condition of typical PM machines [2]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002131_012043-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002131_012043-Figure4-1.png", + "caption": "Fig 4. Schematic of the braking control system for a two-wheeled motorcycle: 1 - front wheel; 2 \u2013 supporting element for the front fork with a lateral force-measuring sensor; 3 \u2013 front wheel RPM sensor; 4 \u2013 cable drive of the front wheel brake; 5 \u2013 front wheel brake lever; 6 \u2013 transmission channels; 7 \u2013 rear wheel brake pedal; 8 \u2013 cable drive of the rear wheel brake; 9 \u2013 front wheel brake; 10 \u2013 front wheel brake disc; 11 \u2013 actuator \u2013 tractive electromagnet of the front brake; 12 \u2013 rear wheel; 13 \u2013 rear wheel brake; 14 \u2013 rear wheel brake disc; 15 \u2013 actuator \u2013 tractive electromagnet of the rear brake; 16 \u2013 bottom supporting element for rear wheel shock absorbers with a lateral forcemeasuring sensor; 17 \u2013 sensor of the actual braking torque of the rear brake; 18 \u2013 sensor of the actual braking torque of the front brake.", + "texts": [], + "surrounding_texts": [ + "1 Wheel slip relative to the road surface can be monitored by identifying the negative sign of the derivative of forces in the wheel-road contact with respect to time. 2 The experimental studies have shown that the negative sign of the force derivative with respect to time is associated with a decrease in the wheel friction coefficients when the wheel contact is slipping in the longitudinal or lateral direction. 3 The algorithm for generating control signals for actuators of wheeled vehicles, which is based on identification of the negative sign of the force derivative in the wheel-road contact, can be used in the active safety system of any wheeled vehicle. 4 The results of tests of braking torque measurement devices show that they can be used to improve brake mechanisms and as sources of information for onboard diagnostic systems for monitoring braking performance of wheeled vehicles." + ] + }, + { + "image_filename": "designv11_80_0003955_ecce44975.2020.9235383-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003955_ecce44975.2020.9235383-Figure7-1.png", + "caption": "Fig. 7. a.) Equivalent circuit model of the proposed suspension; b.) Electrodevoltage connection for the proposed suspension", + "texts": [ + " Noting that the primary goal of each electrode pair is to create a potential difference from the electrodes to the rotor, it can be shown that the innermost electrodes may be combined to one large neutral central electrode\u2014reducing the number of supplies to only 4 per stator. The area of this new neutral electrode is then equal to the sum of each of the three outermost electrodes, as depicted in Fig. 5b However, this leads to a problem not experienced by magnetic reluctance forces: the force from each electrode is coupled with the others. To explain the issue, the equivalent circuit model of the suspension electrodes and conductive rotor of Fig. 7a may be analyzed, in which the voltage for each electrode is referenced to earth ground. In practical electrostatic suspension structures, each electrode has a large enough overlapping area with the rotor compared to the gap length that a uniform electric field may be assumed. This leads to the convention of each electrode forming a parallel plate capacitor with the rotor. With this assumption, we can apply the law of conservation of charge to the system to determine the potential of the disk as a function of the potential on each electrode as follows Vd(C1 + C2 + C3 + C4) = C1Ve1 + C2Ve2 + C3Ve3 + C4Ve2 Vd = C1Ve1 + C2Ve2 + C3Ve3 + C4Ve4 C1 + C2 + C3 + C4 (18) These equations demonstrate that the potential between any electrode and the rotor (and therefore the force) is dependent on the other electrode voltages, i", + " To accomplish this, the voltage of the central electrode, Ve4 may be set as Ve4 = \u2212C1Ve1 + C2Ve2 + C3Ve3 C4 (19) Acknowledging that the central electrode has thrice the area as the outer electrodes and assuming local gap distances are equal, (19) simplifies to Ve4 = \u2212(Ve1 + Ve2 + Ve3)/3. This result allows for decoupling of the outer electrode forces, and by extension, each of the DOFs. However, this also implies that the central electrode voltage is not a control variable. The circuit realization to keep the rotor at ground potential is shown in Fig. 7b With this configuration, the previous equations to control the DOFs (13), (15), (16) are still valid, but a set of transform equations to go from V1, V2, V3 to Ve1, Ve2, Ve3, Ve4 are required. They can be found as follows V1 = Ve1 \u2212 Ve4 V2 = Ve2 \u2212 Ve4 V3 = Ve3 \u2212 Ve4 (20) Inserting the result from (19) and then solving for the electrode control voltages gives Ve1 = 5V1 6 \u2212 V2 6 \u2212 V3 6 Ve2 = \u2212V1 6 + 5V2 6 \u2212 V3 6 Ve3 = \u2212V1 6 \u2212 V2 6 + 5V3 6 Ve4 = \u2212V1 6 \u2212 V2 6 \u2212 V3 6 (21) These expressions allow for continued use of the developed control laws of each DOF and represent a simple way to convert to practical electrode voltages" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001786_s11837-019-03963-1-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001786_s11837-019-03963-1-Figure1-1.png", + "caption": "Fig. 1. The aligned mesostructure features with parallel roads arranged in a square array. The material coordinate system is shown.", + "texts": [ + " This method has been explored to derive effective moduli for two-dimensional (2D) and threedimensional (3D) material systems,18\u201322 with exact solutions obtainable for the case of noninteracting voids (also known as the dilute solution). Specific solutions for ellipsoidal and polygonal voids were presented by Kachanov and Jasiuk.19,20 These methods are applied herein to derive closed-form expressions for the effective in-plane moduli (Exx and Gxy) of the aligned mesostructure. Adopting the terminology from Rodriguez,6 define the aligned mesostructure as one where all of the printed roads are aligned in a square array, as depicted in Fig. 1. The material coordinate system is defined such that the z-axis is aligned with the extruded road. The void shape can be approximated by a four-point hypotrochoid. This approximation allows a simple conformal mapping function to be defined. Then, using the complex variable method of elasticity,17 the stress and displacement fields around a single void in an infinite domain are found, subject to prescribed stresses at infinity. Once the stress and displacement fields are known, the equivalence of strain energy is applied to obtain exact expressions for the in-plane moduli for a dilute concentration of noninteracting voids" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003672_s42835-020-00538-y-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003672_s42835-020-00538-y-Figure6-1.png", + "caption": "Fig. 6 Observation line of the simulation verification", + "texts": [ + " The definite integral is calculated by the Simpson formula [27]. Take the parameters of the stepped PM in Table\u00a02, i.e., h1 12\u00a0mm, h2 24\u00a0mm, r1 7\u00a0mm, r2 11\u00a0mm, and Br 1.2 T. At the same time, the FEM model of the stepped PM with the same parameters is set up by using the local coordinate system of the PM in Ansys software. Draw a magnetic field observation line with a length of 80\u00a0mm along the X-axis, and this line is 0.5\u00a0mm above the PM and 6\u00a0mm from the Z-axis of the PM. The observation line is illustrated in Fig.\u00a06. (8) \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 Bx = Br 4 \u23a7 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23a9 \u222b h 1 0 \u222b 2 0 (z \u2212 z 1 )r 1 cos \ufffd \ufffd x \u2212 r 1 cos \ufffd2 + \ufffd y \u2212 r 1 sin \ufffd2 + \ufffd z \u2212 z 1 \ufffd2 \ufffd 3 2 d dz 1 +\u222b h 2 h 1 \u222b 2 0 (z \u2212 z 2 )r 2 cos \ufffd \ufffd x \u2212 r 2 cos \ufffd2 + \ufffd y \u2212 r 2 sin \ufffd2 + \ufffd z \u2212 z 2 \ufffd2 \ufffd 3 2 d dz 2 \u23ab \u23aa \u23aa \u23aa \u23ac \u23aa \u23aa \u23aa \u23ad By = Br 4 \u23a7 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23a9 \u222b h 1 0 \u222b 2 0 (z \u2212 z 1 )r 1 sin \ufffd \ufffd x \u2212 r 1 cos \ufffd2 + \ufffd y \u2212 r 1 sin \ufffd2 + \ufffd z \u2212 z 1 \ufffd2 \ufffd 3 2 d dz 1 +\u222b h 2 h 1 \u222b 2 0 (z \u2212 z 2 )r 2 sin \ufffd \ufffd x \u2212 r 2 cos \ufffd2 + \ufffd y \u2212 r 2 sin \ufffd2 + \ufffd z \u2212 z 2 \ufffd2 \ufffd 3 2 d dz 2 \u23ab \u23aa \u23aa \u23aa \u23ac \u23aa \u23aa \u23aa \u23ad Bz = Br 4 \u23a7 \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23a9 \u222b h 1 0 \u222b 2 0 \u2212r 1 (x \u2212 r 1 cos ) cos \u2212 r 1 (y \u2212 r 1 sin ) sin \ufffd \ufffd x \u2212 r 1 cos \ufffd2 + \ufffd y \u2212 r 1 sin \ufffd2 + \ufffd z \u2212 z 1 \ufffd2 \ufffd 3 2 d dz 1 +\u222b h 2 h 1 \u222b 2 0 \u2212r 2 (x \u2212 r 2 cos ) cos \u2212 r 2 (y \u2212 r 2 sin ) sin \ufffd \ufffd x \u2212 r 2 cos \ufffd2 + \ufffd y \u2212 r 2 sin \ufffd2 + \ufffd z \u2212 z 2 \ufffd2 \ufffd 3 2 d dz 2 \u23ab \u23aa \u23aa \u23aa \u23ac \u23aa \u23aa \u23aa \u23ad " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001811_j.procir.2019.09.016-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001811_j.procir.2019.09.016-Figure12-1.png", + "caption": "Fig. 12. The tape can fit to the shape of a complex curved geometry", + "texts": [ + " Therefore, it is very hard to have a bend in the model jump from one bended area across a flat region to the next bended area on the real part. In Fig. 10, a tape which was optimized with the evolutionary approach after the preprocessing is shown. It is easy to see that the tape is now able to approximate the curve of the target shape. It is shown, that the combination of an analytic preprocessor with the evolutionary optimization is beneficial. An example of a more complex target shape is shown in Fig. 12. With its curved geometry, the target shape consists of a multitude of triangles. In the pre-process, the triangle centre points are projected into the plane of the trendline. Fig. 11 shows the projection points (blue) and the smoothed SavitzkyGolay filter curve (red). Due to the curved geometry, the averaged normal vector is causing an inclination of the projection plane, which leads to a non-optimal starting solution (Fig. 12a), since the tape is inclined as well. However, the evolutionary algorithm allows a rotation of the tape by varying the rotation angles \ud835\udefc\ud835\udefc\ud835\udc56\ud835\udc56 (Fig. 12b). Thanks to the evolutionary algorithm, the optimization process still develops a near net shape tape geometry. Fig. 9. Optimized tape without using the start bending parameters from the pre-process Fig. 10. Optimized tape using the start bending parameters from the preprocess Fig. 11. The Savitzky-Golay filter curve approaches the projected triangle centre points. From the Savitzky-Golay curve a linear tape curve is derived. The pre-process as well as the evolutionary algorithm have multiple settings which allow a fine tuning of the optimization" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001800_ismsit.2019.8932720-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001800_ismsit.2019.8932720-Figure8-1.png", + "caption": "Fig 8.Experimental setup of the RSEA.", + "texts": [], + "surrounding_texts": [ + "Our RSEA, include a DC motor (2000 rpm, 24 V) with gearbox 1:10, it is used to applied a rotational force. The type of motor driver is (EC-10A). Rotary Encoders sensors (ARC-50B \u2013 10 bit) are included in the system. The load and motor position data are given to the control unit by the encoder sensors. A torque sensor (FUTEK TRS600 50 Nm) mounted in the load of the system to measure the output torque. An inverted pendulum is mounted in the output of the system which can be the model of a robot leg. The control block system comprises an industrial computer (GoogolThech motion controller GT-800 series) and a PWM pulse Generator card (microcontroller ATMEGA128), which is responsible for controlling and filtering RSEA inputs and outputs. Fig.9 illustrates the experimental setup of the RSEA controlled by PID+FF controller embedded in Industrial Computer. Table III present the system parameters." + ] + }, + { + "image_filename": "designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure13-1.png", + "caption": "Fig. 13 CAD of the double arrowhead structure.", + "texts": [ + " The auxetic effect of such a design is obtained by the concurrent out-folding of re-entrant cells and the missing ribs\u2019 rotation mechanism. The double arrowhead design is another variant of the re-entrant honeycomb, which was first founded through the numerical topology optimization method. Based on the actual configuration of the arrowhead, any extension will cause the triangles to expand in the transverse direction while compressions will cause them to collapse. The auxetic behavior of such geometries is shown in Fig. 12. The CAD drawing of the double arrowhead structure is shown in Fig. 13. These structures are designed to exhibit a Poisson\u2019s ratio of 0.8. However, they have been measured to exhibit a Poisson\u2019s ratio of 0.92 at smaller strains (Kolken and Zadpoor, 2017). As mentioned in the previous section, the five designs were all fabricated using the three different materials and assembled into a concentric tube structure with the auxetic material as the inner tube and the normal PPR tube as the outer tube. For each of the assemblies, we performed a threepoint bending test (Instron universal testing machine) as it is the most typical representation of the type of movement that our surgical device undergoes during its course of operation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002950_042044-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002950_042044-Figure3-1.png", + "caption": "Fig 3. Magnetic induction intensity vector diagram.", + "texts": [ + " It is assumed that the permanent magnet and the outer steel frame are stationary, and it is relative to speed between the permanent magnet and the conductor barrel. The output variable is taken as the detection parameter of torque meshing and solved. In this paper, under the condition of 20rpm slip and 40rpm slip, the coupling is simulated by electromagnetics under 50% load and 100% load respectively. Taking 20rpm as an example, the finite element simulation results are introduced. It can be seen from Fig. 3 that the area with the largest magnetic induction intensity in the conductor barrel is the area where the permanent magnet maps to the conductor barrel, and with a maximum value of 1.84t. The magnetic induction intensity between the mapping areas of adjacent permanent magnets drops sharply. The direction of the magnetic induction intensity in the conductor barrel is the same as the magnetization direction of the permanent magnet. The direction of the magnetic induction intensity in the mapping area of adjacent permanent magnets is opposite, and the vector of the magnetic induction intensity is on Through the outer steel plates on both sides, a closed-loop curve is formed with each other" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000353_iicpe.2018.8709523-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000353_iicpe.2018.8709523-Figure1-1.png", + "caption": "Fig. 1. Rolling-element bearing structure and parameters.", + "texts": [ + " DYNAMIC-MODELING OF INDUCTION MOTOR By considering basic assumptions presented in [12]-[16], simplified extended state-space induction motor model based on ,s qdoi ,r qdoi ,r\u03c9 and lT in synchronous reference frame ( )2s sf\u03c9 \u03c0= [16], [17] is formulated as: ( ) ( ) ( ) ( ) 0 0 0 0 0 1 s s s r r s q q s p r d q p r d q s s s r r s d d s p r q d p r q d s s s r s s r r s q q p r d q s p r d q r s s r r s d d p r q d s p r d d r r r em l p i - i - n i i - n i v i - i n i i n i v i v - i i i n i - i - - n i - v i i - n i - i - n i - v i - i T - B n J \u03b1 \u03c9 \u03b2 \u03c9 \u03c7 \u03c2 \u03c9 \u03b7 \u03b1 \u03c9 \u03b2 \u03c9 \u03c7 \u03c2 \u03c9 \u03b7 \u03d1 \u03b6 \u03c3 \u03c4 \u03c9 \u03c5 \u03c9 \u03be \u03c9 \u03c1 \u03c3 \u03c4 \u03c9 \u03c5 \u03c9 \u03be \u03c9 \u03c1 \u03b4 \u03c9 = + + + = + + + + + = = + = + = = ( ) 0 r l l -T T \u03c9 = (1) where, 2 0 0 0 0 2 0 0 0 0 0 0 0 , , , 1 , , , , 3 , , 2 s r m r m m r sr s r m ls ls s m s m r s s r s r sm r em p m q d d q lr r L L r L L L , a a a a rL a L L - L a L L r L L L r L L L , , a a a a L r T n L i i - i i a L \u03b1 \u03b2 \u03c7 \u03c2 \u03b7 \u03d1 \u03b6 \u03c3 \u03c4 \u03c5 \u03be \u03c1 \u03b4 = = = = = = = = = = = = = = = ( )r (2) Here, stator currents are 0i s qd ; rotor currents are 0 r qdi ; rotor speed r\u03c9 ; load torque lT ; and input stator voltage 0.s qdv Healthy induction motor model presented in (1) is used for state estimation. III. VALID BEARING FAULT MODELS A. Bearing Fault Types and their Characterstics Frequencies As most of the induction motors use rolling-element bearings. The schematic representation of rolling-element bearings structure is shown in Fig. 1. The main rollingelement bearing parameters are outer raceways, inner raceways, the balls, and the cage. Bearing faults can be classified into two types: 1) single-point faults, and 2) generalized roughness (permanently damaged bearing) [4]- [7]. Further, single-point bearing defects can be classified as follows: Each type of bearing fault introduces characteristics frequency ( )cf components in the stator current spectrum. The characterstics frequencies are functions of bearing geometry parameters and often related to shaft speed (in turn mechanical frequency)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003468_s10846-020-01251-8-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003468_s10846-020-01251-8-Figure1-1.png", + "caption": "Fig. 1 HL-20 coordinate systems, Euler angles, and body rate angles [41]", + "texts": [ + " The paper is organized as follows. In Section 2, the equation of motion is given and some simplifications for notation are introduced. Observer and controller design and stability analysis of the closed-loop system are given in Section 3. Simulation results are presented in Section 4. Finally, concluding remarks are given in Section 5. In this section, the equation of motion of the HL-20 flight vehicle is derived based on Newton\u2019s law. HL-20 coordinate systems, Euler angles, and body rate angles are shown in Fig. 1. To obtain a decoupled reduced model as mentioned in [6, 42] and deriving the rigid body equation of motion, the following assumptions are made: 1- The body is rigid. 2- There is symmetry in roll aerodynamic. 3- Mass is constant that is dm dt \u22450. 4- The curvature of the earth\u2019s surface can be neglected when considering short distance flights. The rotational rigid body equation of motion is represented as [5, 43, 44]: p\u0307 \u00bc Lpqpq\u2212Lrqrq\u00fe LmxMX \u00fe LmzMZ \u00f01\u00de q\u0307 \u00bc Mrprp\u2212Mr2p2 p2\u2212r2 \u00feMmy MY \u00f02\u00de r\u0307 \u00bc Npqpq\u2212Nrqrq\u00fe NmxMX \u00fe NmzMZ \u00f03\u00de \u03d5\u0307 \u00bc p\u00fe sin \u03d5\u00f0 \u00detan \u03b8\u00f0 \u00de\u00f0 \u00de q\u00fe cos \u03d5\u00f0 \u00detan \u03b8\u00f0 \u00de\u00f0 \u00der \u00f04\u00de \u03b8\u0307 \u00bc cos \u03d5\u00f0 \u00de\u00f0 \u00de q\u2212 sin \u03d5\u00f0 \u00de\u00f0 \u00de r \u00f05\u00de \u03c8\u0307 \u00bc sin \u03d5\u00f0 \u00de=cos \u03b8\u00f0 \u00de\u00f0 \u00de q\u00fe cos \u03d5\u00f0 \u00de=cos \u03b8\u00f0 \u00de\u00f0 \u00de r; \u00f06\u00de where p" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001546_8056342-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001546_8056342-Figure1-1.png", + "caption": "Figure 1: Three-dimensional engagement geometry.", + "texts": [ + " An appropriate coordinate system is conducive to revealing relative-motion characteristics between a missile and a target and simplifying the design of a guidance law. In existing literatures, there are mainly two coordinate systems used to describe intercept engagement. One is the commonly used spherical LOS coordinate system, as shown in [9]. The other is the rotating LOS coordinate system proposed in [5]. In this paper, we adopt the rotating LOS coordinate system where a decoupled relative-motion equation set is obtained. The three-dimensional engagement geometry is presented in Figure 1 where a missile M is intercepting a maneuvering target T . The missile and the target are assumed as point masses. oIxIyIzI represents the inertial reference frame. vm and vt denote the velocities of the missile and the target, respectively. From Figure 1, we have r = rt \u2212 rm = rer , \u00f01\u00de where r is the missile-target relative distance and er is the unit vector along the LOS. Orientation variation of the target relative to the missile brings about rotation of the LOS. However, the LOS rotation is completely dependent on the components of vm and vt normal to the LOS. Therefore, angular velocity of the LOS rotation is perpendicular to the LOS and denoted as \u03c9. Then, \u03c9 is computed as \u03c9 = r \u00d7 v r2 = \u03c9e\u03c9, \u03c9 \u2265 0, \u00f02\u00de where v = vt \u2212 vm, e\u03c9 is the unit vector along \u03c9, and \u03c9, the length of \u03c9, is known as the LOSR" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure6.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure6.3-1.png", + "caption": "Fig. 6.3 Stage 2 of the proposed joining method", + "texts": [ + ", internal tube is larger than the external tube, when axial compression takes place the internal tube starts to deform at the die surface. Since there is no contact between the punch and external tube, it rests at the bottom portion without having any deformation which is shown in Fig. 6.2. 68 E. Premananda and R. Ganesh Narayanan Stage 2: In this stage, the external tube comes in contact with the punch and it also starts deforming above the internal tube. By the time, the internal tube keeps on deforming and acquires the shape of groove of the die. The process of stage 2 is shown in Fig. 6.3. Stage 3: As the upper portion of the die groove is heading inward, the leading edge of internal tube starts deforming in the inward direction. After a particular stage, the tubes get mechanically locked as shown in Fig. 6.4. Some experimental trials were conducted to demonstrate the joining method at lab scale. Figure 6.5 shows the experimental setup clamped in a universal testing machine. The tubes were displaced in the downward direction at a uniform cross-head speed of 1 mm/min. 6 Joining Concentric Tubes by End Forming: A Finite Element \u2026 69 The uniaxial tensile experimental tests were performed in order to study the mechanical properties of the tubes by means of universal testing machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000635_978-3-030-20377-1_1-Figure1.8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000635_978-3-030-20377-1_1-Figure1.8-1.png", + "caption": "Fig. 1.8 Square block on half-space forming a complete contact", + "texts": [ + " Examples include internal corners at, for example, firtree root corners and external corners, for example, in the case of a railway line head where the line head width is only 2\u20133 times the contact size. Because it is not possible to form simple closed-form elasticity solutions to shapes such as rectangles, except by series representation, it is equally impossible to produce simple recipes for the state of stress for complete contacts generally, even thosewhere the bodies have simple shapes, such as that shown inFig. 1.8. It is inevitable, therefore, that numerical methods, such as the finite element method, will have to be used. But, we can obtain at least some useful information in the neighbourhood of the corners of these contacts, and it is these regions which are often of the most practical interest, as it is at or near them that there is the greatest possibility of fretting, or where cracks may start. If we \u2018zoom in\u2019 with a microscope so that the field of view includes just the surfaces near the contact edge, Fig. 1.8, we see something which is simply two wedges\u2014one a half-plane and the other, in this case, a quarter plane. The bodies may be locally separated, or in contact but slipping, or adhered, depending on the loading on the bodies overall, but we shall start off with the simplest assumption\u2014 that they are, at the edge, in intimate contact and adhered, so that they may, together, be thought of as a monolithic three-quarter plane. This is simply a special case of a wedge of included angle 2\u03b1 with free surfaces, and was studied in detail byWilliams in a very celebrated solution which we now present" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure90.13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure90.13-1.png", + "caption": "Fig. 90.13 Buckling of liner during depressurization", + "texts": [ + " additional problem, as during consecutive pressurization-depressurization cycles, it may cause liner to buckle; hence, bringing down the efficiency of liner and possible failure during operation. The eigenvalue linear buckling analysis is conducted to observe the buckling of liner (plastic condition) when a compressive load of 1 MPa is assigned to obtain load at which the liner buckles. During the depressurization phase, if the external load on the liner exceeds the critical buckling load of Fcr = 5.04 MPa the liner will buckle as seen in Fig. 90.13. 90 Finite Element Analysis of Potential Liner Failures \u2026 1085 The three types of failures occurring in liner during operation are analyzed and the following conclusions are drawn: \u2022 The presence of thickness variation in liner forms a transition region. Burst pressure load of 4.32 MPa, applied on liner showed that failure of liner occurred at these transition regions, i.e. equator region (312 MPa Von Mises stress and 0.82% plastic strain) and dome region (358 MPa Von Mises stress and 0.99% plastic strain)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002215_s00170-019-04806-8-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002215_s00170-019-04806-8-Figure9-1.png", + "caption": "Fig. 9 Coordinates of the contact points in the coordinate system Llh1z1 associated to the instantaneous center of velocities of the planar motion", + "texts": [ + " Thus, generalized force Q\u03b3 is composed of moments of assembling forces, gravity force, and friction forces applied at the contact points. Q\u03b3 \u00bc mLl F as \u00fe mLl mg \u00fe mLl FK \u00fe mLl FB1 \u00fe mLl FB2 : The moment of assembling force is a known value that depends on a method of assembly. To determine gravity force mg moment and moments of friction forces FK , FB1, and FB2 relative to instantaneous axis Ll, axes of coordinates \u041e1h1 and Lz1, perpendicular to instantaneous axis of rotation, shall be associated with point L: axis Lh1, parallel to axis\u041e1h, and axis Lz1, parallel to hole axis\u041e1z (see Fig. 9). Gravity force moment (see Fig. 9) relative to instantaneous axis is equal to mLl mg \u00bc \u2212mgHc \u00bc \u2212mg \u041e1L1\u2212\u04211\u041e1\u00f0 \u00de \u00bc \u2212mg S1\u2212hc\u00f0 \u00de \u00bc \u2212mg S1\u2212 0:5H\u2212a2\u00f0 \u00desin\u03b3\u00bd : \u00f024\u00de Moment of each friction force relative to instantaneous axis shall be defined by formula mLl F \u00bc h1Fz1\u2212z1Fh1; \u00f025\u00de where h1 and z1 are coordinates of points for applying these forces in the specified system of coordinates (Fig.8), Fz1 and Fh1 are projections of friction forces to these axes. Coordinates of contact points \u04121, \u04122, and \u041a are equal to h1B1 \u00bc h1B2 \u00bc h1B \u00bc LB3 \u00bc 2S1; h1K \u00bc \u2212KL1 \u00bc \u2212KLcos\u03b3; z1B1 \u00bc z1B2 \u00bc \u2212L1L \u00bc \u22122a1; z1K \u00bc \u2212KLsin\u03b3; \u00f026\u00de Axis of coordinates Lz1 is parallel to axis \u041e1z, hence, direction cosines of friction forces relative to axis Lz1 are equal to direction cosines of angles relative to axis\u041e1z, the values of which are as follows (19) cos\u03bbF B1 \u00bc \u2212 VB1z VB1 ; cos\u03bbF B2 \u00bc \u2212 VB2z VB2 ; cos\u03bbF K \u00bc \u2212 VKz VK : After substituting all components into formula (25) and further transformations, moments of friction forces FK , FB1, and FB2 relative to instantaneous axis of rotation shall be equal to mLl FK \u00bc hK FKz\u2212zK FKh \u00bc f NK KL2\u03b3\u0307 VK ; mLl FB1 \u00bc hB1FB1z\u2212zB1FB1h \u00bc f NB1 BL2\u03b3\u0307\u22122a1b\u03c8\u0307\u22122a2b\u03c6\u0307 VB1 ; mLl FB2 \u00bc hB2FB2z\u2212zB2FB2h \u00bc f NB2 BL2\u03b3\u0307 \u00fe 2a1b\u03c8\u0307\u00fe 2a2b\u03c6\u0307 VB2 : Thus, generalized force of the first Lagrange equation is equal to Q\u03b3 \u00bc mLl F as \u2212mg S1\u2212 0:5H\u2212a2\u00f0 \u00desin\u03b3\u00bd \u00fe f NK KL2\u03b3\u0307 VK \u00fe f NB1 BL2\u03b3\u0307 \u00fe 2a1b\u03c8\u0307\u22122a2b\u03c6\u0307 VB1 \u00fe f NB2 BL2\u03b3\u0307 \u00fe 2a2b\u03c6\u0307\u00fe 2a1b\u03c8\u0307 VB2 : \u00f027:1\u00de Generalized force Q\u03c8 in the second Lagrange equation is equal to the sum of moments of all these forces relative to the hole axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure4.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure4.6-1.png", + "caption": "Fig. 4.6 Topology optimization", + "texts": [ + " It is not enough to say Aluminum has 1/3 of the elastic modulus of steel so we need to beef up the section or AHSS has a 700\u20131500 MPa strength so we can reduce the section from 3 to 2 mm. Such a general statements can be detrimental to design and can also give these advanced materials a bad taste, since the function of the product will not meet the customer expectations. Consider another example to lightweight a steel casting to Aluminum one. The process follows as in Fig. 4.5 with a slight modification to the design, where a topology optimization is needed to substitute for the equation. Figure 4.6 (http:// www.altairatc.com/(S(rfs4o2whprtbiaaazxof4unk))/europe/Presentations_2009/Ses sion_05/SWEREA_Topology%20Optimization%20of%20Castings_091103.pdf) shows an example of such a change in design and topology to yield the desired function. This type of system approach where the design, manufacturing process, optimization and cost structures are considered upfront is really common sense but not practiced regularly in the industry. The problem lies in the education and culture. The educational institutions are becoming silos of information where the theory and virtual tools are not connected and practiced" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000149_robio.2018.8664847-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000149_robio.2018.8664847-Figure2-1.png", + "caption": "Fig. 2. Overview of a cutting liver experiment [4]", + "texts": [], + "surrounding_texts": [ + "In this section, firstly, we describe cutting force measurement of targeted biological tissue, here swine liver is selected as a soft biological tissue, and then we explain the overview of the developed haptic interface using MR fluid ." + ] + }, + { + "image_filename": "designv11_80_0003389_ls.1517-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003389_ls.1517-Figure1-1.png", + "caption": "FIGURE 1 Structural sketch of opposite reciprocating friction and wear tester. 1-Synchronous pulley; 2-opposite eccentric shaft; 3-connecting rod; 4-joint connection mechanism; 5-heating block slider; 6-heating block guide rail; 7-insulation layer; 8-heating block; 9-cylinder liner specimen; 10-piston ring specimen; 11-indenter; 12-roller needle; 13-self-aligning pressure block; 14-resistance strain sensor; 15-plate spring; 16-loading screw; 17-indenter Slider; 18-indenter guide; 19-piezoelectric sensor", + "texts": [ + " These experiments were completed on a self-made cylinder liner-piston ring reciprocating friction and wear tester. The practicability and reliability of the contact resistance method to determine the lubrication state of the cylinder liner-piston ring friction pair were verified under different loading condition. The experimental machine consists of five systems: reciprocating motion, oil supply, heating, loading and resistance signal acquisition. The specific style of the opposite reciprocating friction and wear tester is shown in Figure 1. Its purpose is to simulate the lubrication behaviour of the cylinder linerpiston ring tribological system. The motion form of the friction pair is shown in Figure 2. The piston ring specimen is fixed in the indenter and remains stationary during the experiment. The cylinder liner specimen is fixed in the groove of the heating block. Driven by the connecting rod, the heating block carries the cylinder liner specimen to make linear reciprocating motion with a stroke of 30 mm, and a lubricating oil film is formed between the surface of the piston ring specimen and the cylinder liner specimen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001087_ccdc.2019.8832450-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001087_ccdc.2019.8832450-Figure2-1.png", + "caption": "Fig 2. Steering system model", + "texts": [ + "3 Modeling of Steering System Steering system is a complex mechanism of a vehicle containing many nonlinear factors. To simplify the model, some assumptions are shown below: (1) The rack and pinion steering gear is applied in the towing car. (2) The dry friction and connection gaps between steering linkages are ignored. (3) The left and right wheels swing at the same angle. (4) The deformation of the right and left tie rods is the same. (5) The steering hand wheel is fixed. Based on these assumptions, the three DOFs steering system model can be received (see Fig 2). It is clear that the three DOFs are the rotation of the steering column, the movement of the rack and the rotation of the steering wheel around the king pin. It should be mentioned that the angle between the track arm and the tie rod (denoted by \u03b2) and the angle between tie rod and the rack (denoted by \u03b3) are considered to be approximately constant because the change of steering angel is small. Then the deformation of the steering tie rods can be written as = cos cos( ) cos , 2wl y (8) where denotes the deformation of the tie rod, l denotes the effective length of the steering knuckle arm, is the front caster angle, and y denotes the displacement of the rack" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000613_1.5112677-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000613_1.5112677-Figure3-1.png", + "caption": "FIGURE 3. Initial CAD-Model of the Cowling (a), CAD-Model of tool part (b) and completed tool in STL-Format (c)", + "texts": [ + " Therefore, an indirect manufacturing process is chosen here. The cowling should be made of plastic sheet by means of thermoforming. Overall, only a small series of about 10 cowlings should be formed. The engine bonnet itself has dimensions of approx. 284 x 262x 141 mm. Additive Tooling goes through a process chain that starts with the CAD data and ends with the manufacture of the component (see Fig. 2). The basis for the additive tooling is the CAD data of the cowling, which is provided by the project partner responsible for the aircraft design (see Fig. 3 (a)). These CAD data, which contain only the surface, are first transformed into a solid body that corresponds to the tool. It must also be checked whether there is an undercut. In this case, a slight undercut (less than 1 mm) was found which, due to the flexibility of the formed component, poses no difficulty in demolding. To save printing material, the body is hollowed out, so that a wall thickness of about 30 mm remains (see Fig. 3 (b)). This wall thickness provides sufficient strength for the subsequent forming process. The tool was added by a circumferential edge of about 10 mm height, which is needed for the cutting of the formed part. In addition, a number of channels were placed in the lower part of the tool to allow vacuuming at these critical points during thermoforming. 150001-3 As part of data preparation, the data of the tool are positioned and aligned in the construction chamber of the BJ printer. In addition, the construction parameters (e.g., layer thickness and build speed) can be selected. Since this is a relatively large component, the tool has to be divided into two parts as shown in Fig. 3 (c). After the transfer of the data form preprocessing to the printer, the BJ process can start. In this case a Binder Jetting Printer ProJet 660 from 3D-Systems with a construction camber of 381 x 254 x 203 mm is used. First, from a reservoir an approx. 0.1 mm thick layer of plaster powder is applied to the piston. The excess powder lands in a second reservoir. The print head then moves across the newly applied layer of powder and applies the binding agent and the colour (not needed in this case)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure78.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure78.6-1.png", + "caption": "Fig. 78.6 3D model of from solid works to ANSYS current and modified with hard-facing Fibrizer Hammer", + "texts": [ + "G), angular velocity (x), linear velocity (v), torque (T), Centrifugal force (Fc), Work done in shredding per revolution, Force generated by pin (Fp), Force generated by cane being shredded (Fs) and Stress concentration analysis by which maximum shear stress were calculated for modified Fibrizer hammer. In addition to that, cost comparison was also made between existing and modified Fibrizer hammer. 78 Design Analysis and Modification of Sugarcane Fibrizer Hammer \u2026 939 A three-dimensional model of Fibrizer hammer was developed in a modeling software SOLDWORKS by comprehensively considering the structural feature, processing requirements, and solution calculation amount for current and modified Fibrizer hammers. This model was imported into ANSYS 18.0, a finite element software as shown in Fig. 78.6. The hammer part adopted the structural steel as per the records. Automatic mesh generation was activated and obtained that the number of nodes were 81,074, unit number was 48,701 for current Fibrizer, whereas 42,601 and 24,487 nodes and units were for modified Fibrizer. One end of hammer where circular hole existed was fixed and other end of Fibrizer hammer was given the point load of 7136.6 N as boundary conditions. Load was applied along positive y-axis. Simulation results were recorded on harmonic response, total deformation, von Mises stresses, maximum shear stress, and 940 T" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000157_022097-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000157_022097-Figure2-1.png", + "caption": "Figure 2 Testing bearing", + "texts": [ + " The test rig is mounted on a T-slot cast iron base with a single-disk symmetrical rotor and two identical test bearings placed symmetrically on both ends. An oil supply system is available to provide a steady flow of lubricating oil for the testing system while an oil sink mounted on the T-slot cast iron base can reserve the lubricants. Shaft rotating speed is adjustable range 0-10000 rpm. The parts marked in figure 1 are: 1 speed motor; 2 bearing base; 3 eddy current sensor; 4 shaft; 5 disk; 6 acceleration sensor; 7 PC; 8 measurement and control system The test bearing used in experiment is shown in figure 2. The test bearing is composed of an ordinary sliding bearing and the HSFD. The HSFD studied in this paper is a four oil pad type static pressure squeeze film damper. A restrictor valve is connected to each oil chamber of the damper, and oil is supplied to the damper through the restrictor valve. A restrictor valve is connected to each oil chamber of the damper, and oil is supplied to the damper through the restrictor valve. Parameters of test rig and test bearing are shown in Table 1. CISAT 2018 IOP Conf" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001557_ias.2019.8912014-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001557_ias.2019.8912014-Figure4-1.png", + "caption": "Fig. 4. Top and side view of Joy 10SC32B shuttle car", + "texts": [ + " This controller was chosen primarily because of the wide range of operating modes available, including pulse (RC radio) mode. The controller also permits hall sensor or synchronous serial interface (SSI) rotary shaft encoder signals. The controller and traction motors are supplied by a 22.2 V, 5.4 Ah, lithium polymer (LiPo) battery. A step file of a Joy 10SC32B provided by Komatsu Mining Corp. was used to establish the relevant dimensional ratios and details for the laboratory-scale prototype body. Fig. 4 shows the top and side view of the 3-D drawing file provided by Komatsu. It is noted that some details are not included with the file, e.g., the operator\u2019s platform, cable reel, chain a. Prototype chassis b. Steering servo motors Page 3 of 7 978-1-5386-4539-0/19/$31.00 \u00a9 2019 IEEE 2019-MIC-0825 conveyor, etc. These missing details did not impact the development of the prototype body design. The drawing of the Joy 10SC32B was used to develop a step file of the 1:6 scale body. This model preserved the important dimensional features of the shuttle car, with some simplifications and modifications" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002698_b978-0-12-821918-8.00003-6-Figure3.60-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002698_b978-0-12-821918-8.00003-6-Figure3.60-1.png", + "caption": "Fig. 3.60 A schematic representation of the Kolsky tension bar. (Song, B., Nishida, E., Sanborn, B., Maguire, M., Adams, D., Carroll, J., Wise, J., Reedlunn, B., Bishop, J., Palmer, T., 2017. J. Dyn. Behav. Mater. 3, 412. With kind permission of Springer Nature.)", + "texts": [ + "36 represent the front and back faces, respectively. The engineering strain rates and the engineering strain in the specimens are evaluated by _\u03b5\u00bc 2C0 Ls \u03b5r (3.37) \u03b5\u00bc 2C0 Ls \u00f0t 0 \u03b5rdt (3.38) 111Testing: Comparison of AM data with traditionally fabricated whereC0 is the elastic stress wave speed in the bar material and Ls is the gage length of the specimen. Note that a Kolsky bar is an apparatus for testing the dynamic stress-strain response of materials. A schematic illustration of the Klosky compression bar is seen in Fig. 3.60. Further to Fig. 3.56, a comparison between the dynamic compressive stress-strain curves at various strain rates of the wrought and the AM 112 Additive and traditionally manufactured components 304L specimen in the Z direction are seen in Fig. 3.61. As can be seen, the two groups are assembled, each in its own fabrication method. The elasticplastic characteristics of the two groups are quite similar. The thick lines represent longitudinal orientation and the thin lines transverse specimen orientations (directions)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003972_icma49215.2020.9233579-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003972_icma49215.2020.9233579-Figure1-1.png", + "caption": "Fig. 1 Concept of the robotic wheel for locomotion and climbing over an obstacle (top), pictures of an assembled robotic wheel (bottom).", + "texts": [ + " In t r o d u c t io n There are generally two types of robotic locomotion introduced for the robust mobility on different types of grounds. One is to use its legs and the other is to employ its wheel. The former is suitable for the efficient locomotion on a rough ground, and the latter gives better performance on a flat ground and a slope. In this research, we pay attention to the moving ability of a wheel, and aim to introduce the new functional mobility of the robotic wheel locomoting on a rough terrain. Fig. 1 shows the concept of climbing over obstacle together with the picture of our robotic wheel. Different robots with a specific wheel to climb the obstacles have been introduced so far. For example, Tadakuma et al. developed a robot with spherical omni wheels that can move all directions and even on uneven terrains [1]. Herbert et al. designed a robotic wheel shaped like a boomerang [2]. This robot has high locomoting ability on rough terrains without having variable wheels. Alsalman et al. simulated a wheel that could locomote on rough terrain without deforming its wheel actively but had a passive mechanism [3]", + " The minimum required torque is estimated to be 5.2kgf-cm, and the servo motor is selected to meet the requirement. The specification of the servo motor (ROBOT SERVO RS301CRF3 FUTABA) is shown in Table 1. 1) Mechanical Design of the Wheel We design the robotic wheel by using a CAD software. The design and appearance of the entire body is shown in Fig. 2. The foundation is designed with a rail structure that allows the spokes to slide stably as shown in Fig. 3. The picture of the assembled robotic wheel based on the design drawing is shown in Fig. 1. The robotic parts are printed by a 3D printer (AGILISTA-3200, KEYENCE) and also fabricated by a modeling plotter (NC-5SK, Mimaki). The materials of the parts are acrylate resin and ABS. 2) Rack & Pinion Mechanism A pinion gear connected with a servo motor is combined with a rack gear with a spoke as shown in Fig. 4. The extension can be achieved by the rack & pinion mechanism up to the half length of a radius. Each spoke moves in the direction of extension from the initial radius of the wheel and 1258 Authorized licensed use limited to: San Francisco State Univ" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001669_icems.2019.8921899-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001669_icems.2019.8921899-Figure4-1.png", + "caption": "Fig. 4. Leakage magnetic field distribution of stator end windingss.", + "texts": [ + " STATOR BARS The distribution of the magnetic field of stator end region in yoz plane under rated condition at 40ms is shown in Fig. 3. The magnetic field of stator end region is generated by the rotor field winding current and the stator end winding current. Eddy currents in clamping fingers and clamping plates influence the distribution of the end leakage magnetic field. In addition, the magnetic shield can absorb some end leakage magnetic flux. The distribution of leakage magnetic flux density for the end of stator winding under rated condition at 40ms is shown in Fig. 4. The end leakage flux density is larger for the stator bar with larger current at a certain time under rated condition. The top layer bar has larger leakage magnetic flux density than the bottom layer bar for the same stator slot in the line segment of stator bars. The leakage magnetic flux density distributes very complicated in the involute segment of stator bars, which is not only influenced by the current of the stator bar itself, but also influenced by the current of adjacent stator bars. As a whole, the leakage magnetic flux density of the stator bar decreases gradually along the direction far away from the stator iron core in the involute segment of stator bars" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002052_b978-0-12-812750-6.00001-9-Figure1.16-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002052_b978-0-12-812750-6.00001-9-Figure1.16-1.png", + "caption": "FIGURE 1.16 Calculation of the average height of force that a vehicle imparts on a barrier.", + "texts": [ + " The ratio between the areas subtended by the curve in the compression and restitution phases, under square root, is equal to the coefficient of restitution: \u03b55 ffiffiffiffiffi Er Ea r \u00f01:11\u00de The instrumented barriers allow measuring not only the resulting force but also the distribution of forces on the impact surface during the impact. Fig. 1.15 shows an instrumented barrier, composed of many load cells arranged on the impact surface. Summing all the load cell values, the resultant force is obtained. The aver- age height of force that a vehicle imparts on a barrier is shown in Fig. 1.16. During the impact, the average height of force is generally changing, as illustrated in Fig. 1.17 for two types of vehicle, a family car and an SUV. More generally, during the impact, the distribution of forces changes. In Figs. 1.18 and 1.19 the distribution of forces for an SUV at 34 and 56 ms from the beginning of the impact is shown. The peak of the force is in correspondence with the rails. In Fig. 1.20, on the other hand, the distributions of the impact forces are shown for two different vehicles, a family car and an SUV" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure7-1.png", + "caption": "Fig. 7 Re-entrant structure CAD.", + "texts": [ + " However, care should be taken, as slight deviations from this range of temperature can quickly make the material brittle and hard. Pythonflex: Pythonflex is a high-performance variant made from specially formulated TPU material. As mentioned earlier, we used five different auxetic designs to cover a wide range of concepts of auxetic behavior and to explore the robustness of our results. Each of our five designs were printed with the three materials mentioned in the previous section and subjected to mechanical tests. The conceptual theory behind the auxetic behavior of each design is elaborated below. Fig. 7 shows a traditional hexagonal re-entrant structure. When a force is applied in either direction, the diagonal ribs move and rotate in a way to produce an auxetic effect in the other direction. The auxetic effect is observed as the diagonal ribs aligned along the horizontal direction move apart in the vertical direction under tension. Tests have shown that most of the structures involving re-entrant honeycombs undergo deformation. A chiral formation is defined as a nonsuperimposable mirror image" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003742_icarm49381.2020.9195386-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003742_icarm49381.2020.9195386-Figure8-1.png", + "caption": "Fig. 8. Hardware system of the orthois.", + "texts": [ + "org/projects/c3dtoolbox; the content has disappeared since Dec 15, 2019. A Proportion Integration Differentiation (PID) controller is used to drive the ankle and foot complex to follow the reference angular curve. Fig. 7 shows the results of reference angular curve tracking. It shows that the actual angular curve is approximate to the reference angular curve. Therefore, it can be concluded that, in real condition, the orthosis has the potential to help the subject to obtain the normal gait sytle. The hardware system of the orthotic device is illustrated in Fig. 8. Specifically, it is mainly composed of the orthosis, the power supplier, data transmission module and the sensing and control units. The power supplier is a DC Li battery, 3000 mAH and 25 V; it provides the power to all electrical equipments in the system. The data transmission unit is a bluetooth module, HC 08; it transmits the gait data, such as the angle of Link 1, to the computer or smartphone wirelessly in real time. The control unit includes DC electric motor (MAXON RE35), motor driver (MAXON ESCON 50/5) and microcontroller (Arduino UNO)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002609_dese.2019.00022-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002609_dese.2019.00022-Figure2-1.png", + "caption": "Fig. 2. Differential drive and steering mechanisms. a) driving structure. b) Lims yawing structure.", + "texts": [ + " The limbs dynamic motion model includes the kinetic and potential energy interaction model, which approaches a Euler-Lagrange solution. The presented approach provides simulations that concern the underactuated mobility space of a hexapod. In section II, the proposed underactuated mechanical design and the Klann limb kinematics are presented. In section III presentes the deduction of the dynamic control walking model and section IV provides the conclusion. The present work proposes a locomotion mechanism consisting of a differential driving system (Figure 2a), an all-limb synchronous bidirectional steering yaw mechanism (Figure 2b) and the Klann-based limbs (Figure 3). The right and left sided are differential speeds providing instantaneous velocity vt and yileding instantaneous yaw speed, \u03c9t. One drive per lateral triplet of limbs, interconnected front/back sides by tracks from the central drives. The steering mechanical system of Figure 2b shows bilateral direction angle for all-limb synchronously and can turn in yaw \u2212\u03c0/4, ..., \u03c0/4. The Klann mechanism (Figure 3) is an underactuated planar multi-link system, which from rotary input motion, it produces as output a cycloid trajectory (Figure 4). The links proportions are defined to optimize linear motion of the contact point at every rotary half cycle of the crank. The contact point lifts during the other rotary half cycle, before returning to the staring position. Klann [9] presented his famous invention, the \u201dwalking device\u201d, which is usually deployed to emulate biological limbs motion, mostly antropods" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000293_012101-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000293_012101-Figure3-1.png", + "caption": "Figure 3. Maximum shear stress of (a) structural steel, (b) alloy steel, and (c) carbon steel.", + "texts": [ + " The results of Maximum principal stress showed that alloy steel had 407.33 MPa, Structural steel had 411.37 MPa and Carbon steel had 408.74 MPa. The principal stress coefficient plays an important role in increasing strength [13]. This causes consequences on research\u2019s result, proven that structural steel has the highest strength because of the high value of principal stress. IC2MAM 2018 IOP Conf. Series: Materials Science and Engineering 515 (2019) 012101 IOP Publishing doi:10.1088/1757-899X/515/1/012101 Maximum Shear obtained specifically for vertical stresses. Figure 3 shows maximum shear stress for three materials, which is structural steel 481.03 MPa, alloy steel 485.11 MPa and carbon steel 482.44 MPa. Shear stress only shows the hardening of the matrix [14]. The highest shear stress value shows high hardening, and this applies to alloy steel so, alloy steel\u2019s strength is the lowest. IC2MAM 2018 IOP Conf. Series: Materials Science and Engineering 515 (2019) 012101 IOP Publishing doi:10.1088/1757-899X/515/1/012101 The analysis results showed that Structural steel had the equivalent stress, maximum principal stress and maximum shear stress of 921" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003824_aiea51086.2020.00073-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003824_aiea51086.2020.00073-Figure2-1.png", + "caption": "Figure 2 Comparison of different types of J components", + "texts": [ + " The combination of components with different specifications is suitable for the connection of wires with different diameters[6]. The J-type clamp has the characteristics of multiple series and models, and can be adapted to various cable connection requirements in actual use. At the same time, it is often difficult to use a unified installation tool for the installation of different types of J-type clamps. Take the B series of J-type clamps as an example. This series has three models of B1, B2 and B3, as shown in Figure 2. Among them, B3 has one wire slot, and B1 and B2 have two wire slots, labeled L and S, respectively. Among them, the aperture of the L-trough is slightly larger than that of the S-trough for accommodating thicker wires, and S is for accommodating thinner wires[7]. . It can be seen that different types of J-type clamp components have different wire slot positions and different sizes. The traditional clamping method can be applied to a certain wire groove of a certain model, but it cannot support L and S wire grooves at the same time, nor can it be well adapted to different wire diameters" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001713_pierl19060309-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001713_pierl19060309-Figure4-1.png", + "caption": "Figure 4. Contrastive analysis of position effects. (a) Two-dimensional distribution of magnetic flux density. (b) One-dimensional curve of magnetic flux density.", + "texts": [ + " The simulation is simplified to a two-dimensional field, and the magnetic field distribution around the magnetic nanoparticles cluster is simulated based on the uniform medium and static magnetic field distribution. The simulation results are shown in Fig. 3. However, due to the limitation of the simulation model, there is not enough distance between the three clusters of magnetic nanoparticles. Besides, the influence of magnetic field on each other cannot be neglected. Therefore, the results in Fig. 3 may have errors caused by such reasons. In order to ensure the accuracy of the results in Fig. 3 and not affected by the factors of alignment position, the comparative simulation as shown in Fig. 4 shows that different alignment positions will affect the magnitude of magnetic field around MNPs. In addition, the regions with high magnetic field intensity mainly exist near the left and right surfaces of magnetic nanoparticles clusters. With the increase of the radius of magnetic nanoparticles cluster, the range of action in the strong magnetic field region increases. With the decrease of the radius of magnetic nanoparticles cluster, the flux density decays faster, and the magnetic field gradient increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002451_rcar47638.2019.9044101-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002451_rcar47638.2019.9044101-Figure2-1.png", + "caption": "Fig. 2. The communication topology gragh with four nodes and the desired foemation shape.", + "texts": [ + " Coordinating control This paper separates the coordinating control part from a single quadrotor control part, which can reduce coupling of multiple quadrotors control. In first level of the proposed control architecture, a coordinating control algorithm will be designed to get the virtual position and the virtual velocity, which can enable the virtual velocity consensus and the virtual formation shape to be reached in the presence of a virtual leader. cPt and ot can be obtained as follows: Quadrotori dimensional space. From the Fig. 2, we can get the adjacency matrix A. A ~ [~ !~ ~] (22) After the adjacency matrix A obtained, the Laplacian matrix L(A) = [iij] E jRnxn can also be obtained. For every quadrotor, its parameters we used are shown in Table I, and the significance they represented has been explained in the section of dynamic model. In addition, the air resistence coefficient kx,i = ky,i = kz,i = 0.02 and the torque coefficient kq\"i = ke,i = k,p,i = 0.1. The parameters of simulations have been listed in table IT, which are applied to the coordinating cotrol algorithm and the tracking control algorithm", + " And combining the formula (2) and (15), the square of four rotors' speed can be obtained. Theorem 1: Consider coordinating control algorithm (6) and (7). if the virtual leader has a directed path to some of followers and communication topology G = {V, E, A} has undirected path among followers at any time, it can be obtained that Vi -+ Vo and f i -+ f o in finite time, iE~. Proof: The proof can be seen in the Journal paper. In this paper, quadrotors employ the dynamic model of (3) and (4), whose communication topology graph is shown in Fig. 2. This communication topology graph also decides the formation shape of multiple quadrotors. The undirected edges among three quadrotors are discribed by the adjacency matrix A, and the weights of the edges are represented by D = [dij ], which denotes the distance of every two quadrotors in three Based on the proposed control architecture, the trajectory of quadrotors and the virtual leader in three-dimensional space are shown in the Fig. 3, where three quadrotors fly in expected for mation geometric pattern and reach velocity consensus, rapidly" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001713_pierl19060309-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001713_pierl19060309-Figure3-1.png", + "caption": "Figure 3. Flux density distribution of MNPs with different radii. (a) Flux density two-dimensional distribution. (b) Flux density one-dimensional curve.", + "texts": [ + " Based on this principle, a magnetic nanoparticle cluster with radius of 50 (SA-BiotinModifiedFe3O4 Microspheres by Beijing Zhongkeleiming Daojin Technology Co, Ltd), 70 (PLL coated Fe2O3 nanoparticles) and 90 nm (SABiotinModifiedFe3O4 Microspheres) is placed in a uniform background magnetic field with magnetic field intensity of 4 \u00d7 105 A/m. The central positions of the three clusters are (0, 300), (0, 0), (0, \u2212300) (nm), respectively, and their relative permeability is set to 6. The simulation is simplified to a two-dimensional field, and the magnetic field distribution around the magnetic nanoparticles cluster is simulated based on the uniform medium and static magnetic field distribution. The simulation results are shown in Fig. 3. However, due to the limitation of the simulation model, there is not enough distance between the three clusters of magnetic nanoparticles. Besides, the influence of magnetic field on each other cannot be neglected. Therefore, the results in Fig. 3 may have errors caused by such reasons. In order to ensure the accuracy of the results in Fig. 3 and not affected by the factors of alignment position, the comparative simulation as shown in Fig. 4 shows that different alignment positions will affect the magnitude of magnetic field around MNPs. In addition, the regions with high magnetic field intensity mainly exist near the left and right surfaces of magnetic nanoparticles clusters. With the increase of the radius of magnetic nanoparticles cluster, the range of action in the strong magnetic field region increases. With the decrease of the radius of magnetic nanoparticles cluster, the flux density decays faster, and the magnetic field gradient increases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000440_icmech.2019.8722850-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000440_icmech.2019.8722850-Figure2-1.png", + "caption": "Fig. 2. Ball-on-plate experimental setup", + "texts": [ + "40-K1-B00 with i = 26 planetary gearbox) \u2022 2 motor drive controllers (miControl mcDSA-E40-HC) \u2022 2 rotary encoders (SICK DFS60A-BHPA65536) \u2022 a camera (IDS UI-3160CP) \u2022 a microcontroller (MCU, Texas Instruments TMS320 F28335) \u2022 a personal computer (PC) Communication between the microcontroller and the drive controllers is performed via CAN bus using the CANopen protocol. The rotary encoders are connected to the MCU directly using the quadrature encoder pulse (QEP) interface. The PC is connected to the ball-on-plate system using a second CAN bus, where a custom protocol is used. Due to the large size of the plate, a design goal of the actuation mechanism was to have no components horizontally beside the plate. Revolute joints with perpendicular axes in serial configuration, shown in Fig. 2, were placed below the plate to achieve a small footprint. The first joint connects the base frame to an intermediary frame, the second joint connects the intermediary frame to the bracket holding the plate. This mechanism requires a vertical offset between the axes of rotation and the plate. The system acts as an inverse pendulum because the center of mass (CoM) of the plate and bracket assembly is above the axes of the joints. Because the plate surface is above the joint axes, tilting the plate in one direction will cause the ball to move up the slope initially before it is accelerated downwards by gravity, this nonminimum phase behavior must be considered when designing a controller" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000933_s10010-019-00352-7-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000933_s10010-019-00352-7-Figure5-1.png", + "caption": "Fig. 5 Gear processing error evaluation analysis 3", + "texts": [ + " According to gear meshing principle, the tracking error of the X, Y, C axes influences the tooth profile deviation, as shown in the above pictures, and the tooth profile deviation is caused by the motion component of Ex in the meshing line, hence the motion component Ex 0 can be calculated as the following: Ex 0 = Exsin \u02db (2) Fig. 4 illustrates that Ey , Ea, !C , !B and are the tracking error of the Y-axis, tracking error of the A-axis, rotational speed of the workpiece axis, rotational speed of the hob axis, setting angle of the hob, respectively. Furthermore, Ey 0 is the motion component of Ey in the theoretical axis of the hob and can be deduced as: Ey 0 = Ey cos Ea (3) Decomposing Ey 0 to the transverse plane of the work- piece, Ey 00 can be derived as Ey 00 = Ey 0cos = Ey cos Ea cos (4) As shown in Fig. 5, according to the movement of the Y axis, Ey 000 is the motion component of Ey 00 in the meshing line which generated tooth profile deviation, and Ey 000 can be calculated as below: Ey 000 = Ey 00cos \u02db = Ey cos Ea cos cos \u02db (5) As shown in Fig. 6, Ec is the tracking error of the workpiece axis, Ec 0 is the specific error of the workpiece, RC is the workpiece pitch circle radius, !C is rotational speed of the workpiece axis and \u02db is the pressure angle, based on the geometric correlations in Fig. 6, Ec 0 can be derived as: Ec 0 = RC Ec = mt ZC 360 Ec = mnZC 360 cos \u02c7 Ec (6) wheremt , mn, \u02c7, ZC are the workpiece transverse module, workpiece normal module, helix angle and workpiece tooth number, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001552_s12555-019-0234-y-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001552_s12555-019-0234-y-Figure9-1.png", + "caption": "Fig. 9. Kinematic suturing model with the proposed Y-PY-P-Y bending joint.", + "texts": [ + " -PY stacking sequence with 2N + 1 units without having to perform complex calculations. 4. RESULTANT DISTAL ROTATION The application of a CV joint to the wrist of a laparoscopic instrument allows for complete conversion of proximal axial rotation to distal rotation. Distal rotation is an essential movement for suturing an incision with a Cshaped curved needle. With use of the proposed 2-DOF discrete bending joint using 2N + 1 units, a laparoscopic instrument can perform the suturing motion successfully with a curved needle. Fig. 9 shows the kinematic model of a laparoscopic instrument with a Y-P-Y-P-Y bending joint for suturing. Table 5 summarizes the D\u2013H parameters of this model, where the last row can be ignored in many cases. The second row represents Table 2, that is, the transformation of the bending joint. If a desirable joint configuration is applied, this part can be replaced with Table 3 to simplify the calculation. In this case, almost pure distal rotation occurs when the condition \u03b8R +\u03b1B = 0 is satisfied, as depicted in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002050_iros40897.2019.8968099-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002050_iros40897.2019.8968099-Figure1-1.png", + "caption": "Fig. 1. Quadcopter in + configuration.", + "texts": [ + " In addition, the term I\u03c9B denotes that \u03c9 belongs to B and is expressed in frame I . Angular velocity vector of the vehicle is represented by \u03c9B = (p, q, r)T where p, q and r are roll, pitch and yaw rates, respectively. Finally, \u2016\u03c9\u2016 represents the 2-Norm of the vector \u03c9 and |s| represents the absolute value of scalar s. In this section a complete dynamic model of the quadcopter is given, taking the aerodynamic model of the propellers\u2019 thrust force and moments in presence of freestream into account. Figure 1 shows the schematic of a quadcopter. Six reference frames are defined, one of which is assumed to be fixed and attached to the earth, also known as inertial frame I , one is attached to the center of mass of the vehicle and is represented by B and four other reference frames attached to the center of mass of the ith motor Mi, however they do not turn with the rotors. Each propeller generates a thrust force fi in the direction of z-axis of the motor frame. Propellers 1 and 3 have negative and propellers 2 and 4 have positive angular velocities expressed in the body frame as \u03c9pi = (0, 0, \u03c9pi)T ", + " In the end, a discussion on the results and their significance is provided. Generally, in multi-rotor UAVs, hovering is defined as: maintaining a position with zero angular and linear velocities. However, in case of one rotor failure in a quadcopter and in order to control the attitude and altitude of the vehicle, a new hovering definition would be required as: maintaining an altitude while rotating with constant angular velocity about a unit vector that is fixed with respect to the vehicle [1]. Suppose motor number 4 (see Fig. 1) failed. Because of the unbalanced moments of the remaining functioning propellers, the vehicle starts rotating about a unit vector n (as expressed in the body frame) with angular velocity \u03c9B . The evolution of this unit vector in time can be written as follows: n\u0307 = \u2212\u03c9B \u00d7n. (5) According to this new hovering definition, we attempt to keep the orientation of this unit vector fixed with respect to the vehicle. If this unit vector is fixed, from (5) one can conclude that the angular velocity of the vehicle will remain parallel to this unit vector so the vehicle will be rotating about n", + " We introduce a tilting angle \u03b1i about the x-axis of the motor frame Mi to the rotors similar to that in [3] as shown in Fig. 4. According to [3], having \u03b11,3 > 0 and \u03b12,4 < 0 adds to the passive stability of the vehicle in yaw motion which is an improvement in quadcopter flight without any rotor failure. But we are interested in finding the effects of this tilting angle on the mechanical power of the quadcopter after rotor failure and also on spinning UAVs like bispinner. A new configuration is proposed by tilting the rotors about the x-axis of the motor frame (shown in blue in Fig. 1) as shown in Fig. 4 (a) where the positive direction of the tilting angle \u03b1i is shown in Fig. 4 (b). Because rotors 1 and 3 are assumed to be turning in the negative direction of z-axis of the body frame, by tilting these motors by any positive angle, the vehicle tends to generate a yaw motion that is in favor of reducing mechanical power (11). Whereas for rotors 2 and 4 which are turning in the positive direction of the z-axis of the body frame, the tilting angle should be negative. Note that, for simplicity, it is assumed |\u03b11| = |\u03b12| = |\u03b13| = |\u03b14|" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure6-1.png", + "caption": "Fig. 6. Two steps slotting model (Proposed Model)", + "texts": [], + "surrounding_texts": [ + "Using the FEMM 4.2 and coupled with LUA 4.0 scripting, the PMMs characteristics were investigated [7],[11],[12]-[14]. By the combination of FEMM and LUA 4.0 could increase a quick execution for the implementation of a complete simulation of a specific PMMs. Another advantage of LUA application is the capability of parallel computation can be achieved. At the beginning of each simulation, the simulated of PMMs structure were generated in Auto-CAD then exported to the FEMM file. Comparisons of air gap magnetic flux distribution and CT for the PMMs studied were investigated. This means that the proposed PMM model (twosteps slot) was promising for the presence of two steps of slotting in the magnets that do not distort the balance of magnetic force in the air gap of the machine of one-step model. Figures 4, 5 and 6, shows that the value of the flux distribution due to the changing in the magnet structure will not destroy the value of the machine's core losses, as it could be observed that all the value of the flux density approximately 1.436 Tesla. It has been found that if it implemented to the proposed method, it will not change the core losses of the proposed model." + ] + }, + { + "image_filename": "designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.17-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000749_b978-0-12-816716-8.00007-3-Figure7.17-1.png", + "caption": "FIGURE 7.17", + "texts": [ + " (4) The droplet is pushed on the Contact angle of a patterned surface with a varying groove width. textured surface layer by layer, mainly because after the droplet is blown apart, the small droplet enters and fills the groove area, further enhancing the adsorption capacity of the textured surface to the droplet [9]. On an absolutely smooth, flat surface, the liquid presents a Young contact angle with a unique value. However, surface roughness and heterogeneity exist [10], and it is possible that the droplet presents a substable state on the surface, namely the contact angle hysteresis. Fig. 7.17 shows the apparent contact angle hysteresis of liquid drops on the inclined solid surface. As can be seen from the figure, the contact angle \u03b8a in front of droplet slip is significantly greater than the contact angle \u03b8r on the tail side. On an absolutely smooth surface \u03b8a5 \u03b8r, the droplets begin to roll as soon as the surface tilts. But because of the surface heterogeneity \u03b8a 6\u00bc\u03b8r, the solid must be tilted at an angle to allow the drops to tumble. This phenomenon is called contact angle hysteresis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001319_memsys.2019.8870648-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001319_memsys.2019.8870648-Figure1-1.png", + "caption": "Figure 1: Lens dimensions comparison Top: Varioptic 39N0 (15 mm outer diameter, 3.9 mm aperture), Optotune EL-6-18 (18 mm, 6 mm), Optotune EL-10-30 (30 mm, 10 mm). Bottom: Our lens, un-filled (9 mm, 7.6 mm) and packaged. (10 mm, 7.6 mm).", + "texts": [ + " In either case, a fluid acts as a refractive medium, in some cases also an elastomer [3] or a gel [1]. Alternatively, electrowetting is used to create a lens surface between two fluids [6] or a lens effect may be created with liquid crystals [7][8] or acoustic gradient fields [9][10]; for a more complete review see, e.g., [11]. While the latter operate only in resonance at a few 100 kHz, the others have \u22652 ms response time. All have typically half or less of the device diameter optically usable, see fig. 1. In [2], we introduced a novel piezo-glass-piezo sandwich membrane shown in fig. 2 where two piezo rings directly deform an ultra-thin glass membrane. This is then mounted on top of a passive fluid chamber, see fig. 3. We achieved a response time of 2-3 ms and an outer diameter of 19.4 mm with a 12 mm free membrane. This set-up was mostly limited by the elastic support ring that was needed to account for the fluid displacement by allowing an over vertical displacement of the lens membrane. We extended this principle in [3] to an elastomer refractive medium, achieving 79% clear aperture (69% packaged)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003864_ies50839.2020.9231879-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003864_ies50839.2020.9231879-Figure2-1.png", + "caption": "Figure. 2. Overall overview of our system. When the robot walking on uneven surface, the robot performs balancing by adjust body posture to the level of surface flatness", + "texts": [ + " Humans will make adjustments to the tilt and also the speed of the step when walking on a surface that is not ideal. We will combine the inverted pendulum control with the adaptive walking trajectory system, to get the robot's stability when running on the uneven surface. From the results of some of these studies, EROS robots have been able to walk quite robustly on even synthetic grass. Experiments show encouraging results, this method has been successfully implemented in the real robot and successfully achieved remarkable results in soccer games. II. M e t h o d o l o g y The overall system is shown in Figure 2. For a robot to be able to run stably on an uneven surface. So in this research the walking movement generation method is applied based on corrections from the Center of Pressure (CoP). Before starting the walking movement, the initial standing posture of the robot or the robot's default position will greatly affect the balance of the robot when moving. To find out whether the robot's posture has been balanced or not, a center of pressure (CoP) reading is done on the sole of the robot's feet. Corrections are then made if the CoP is not in its proper position" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002253_1350650120908116-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002253_1350650120908116-Figure15-1.png", + "caption": "Figure 15. Load curves on different pump operation conditions with the gear rotation position.", + "texts": [ + " The pressure caused by oil trap is further simulated at the monitoring point in Figure 14. It can be seen that the curves seriously fluctuate with a high amplitude, which is due to the abnormal compression of oil by gear teeth mesh. High output pressure substantially increases the amount of load fluctuation to 147.6%, 230.7%, 341.5% in amplitude compared with the lowest case, which leads to a more serious load on the shaft. Note that the fluctuation frequency tends to remain nearly the same regardless of the pump output pressure. Figure 15 displays the radial load to the gear shaft for single gear meshing cycle. As the geometry of each gear teeth is similar, the load curves repeat periodically with the rotation of gear shaft and there are 10 peaks associated with the teeth number. When the output pressure increases, the average value of the load has an increasing trend of 32.1%, 90.4% (x direction) and 31.5%, 80.1% (y direction). The gear positions (A, B, C, D) are marked in the gear structure diagram in Figure 15, and every point can also be found in load curves. Because the load result has a strong correlation with the CFD flow simulation, it contains a large amount of information of gear pump characteristics including internal fluid, oil trapped, gear meshing, and so on. Analysis of bearings on different pump operating conditions Figure 16(a) shows the locus of the shaft center with respect to the pump load effects during the startup. For benchmark purposes, we present the results on a constant load case. The trajectory shape is seen to be slightly different, and the difference in the ending of trajectory (hydrodynamic-lubrication state) is more obvious than the beginning (mixed-lubrication state) of that. As revealed in Figure 15, the load curves in X and Y directions have a fluctuation due to the transient field variation and it is the reason for the irregular shaft motion. In addition, the average Reynolds number with respect to the radial clearance is less than 1000, consistent with the use of the laminar flow. The variations of the hydrodynamic bearing force and asperity contact force are described in Figure 16(c) for the bearing operating at different load conditions. The different type forces having load-supporting effects are directly related to the pump load fluctuation: the curves with pump load vary around the constant load one" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003110_s40194-020-00955-7-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003110_s40194-020-00955-7-Figure4-1.png", + "caption": "Fig. 4 Experimental verification for (a) powder flux distribution and (b) powder transport ratio", + "texts": [ + " The spatial distribution of powder particles from the lateral nozzle in the experiment was close to Gaussian, which had been verified [8]. According to the modeling approach, the experimental verification for the particle concentration distribution after injection parameters adjustment was mainly from the perspective of key point coordinate information, such as outflow position (A, A1, A2), injection position (O, M, N), and boundary of powder flux (P1, P2), which would be measured by a coordinate measuring machine (CMM). This entire experimental process is shown in Fig. 4(a), and the scheme is listed in Table 3. Since the CMM has an independent measuring platform, a set of experimental devices was designed to approximate the powder feeding in real laser cladding process. The defocusing distance was controlled by a height-adjusting knob and vernier caliper, and the powder feeding angle was controlled by an angle-adjusting knob and inclinometer, which could be magnetically attracted to the powder feeding tube. After each powder feeding condition parameter was determined, powders sprayed from the nozzle were projected onto the surface of an A4 sheet of paper with glue, and then fixed by a transparent PVC plate. Each key point position could be measured by CMM after marking. The experimental scheme was designed to first raise the defocusing distance of the lateral powder feeder from the initial position by 4 mm, and then rotate the feeding nozzle by 3\u00b0 clockwise. A total of three groups of measurements were developed, and the data of 12 measurement points are listed in Table 3. In Fig. 4(b), a powder collection device equipped with two steel sheets, springs, and hexagon socket bolts was used to verify the powder transport ratio under different powder injection parameters. All tests were programmed with 2.322 g/min powder feed rate during 10 s and the placement was marked to ensure that each test has the same initial state before parameter adjustment. As the influence of molten pool width had been developed and compared with theoretical values in Fig. 7 of [8], powder transport ratio had been measured only after adjusting the defocusing distance and powder feeding angle respectively, which ranged from \u2212 2 to 10 mm by 2 mm increments for \u03c7, and from \u2212 3\u00b0 to 3\u00b0 by 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003112_j.mechmachtheory.2020.103903-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003112_j.mechmachtheory.2020.103903-Figure5-1.png", + "caption": "Fig. 5. Fully parallel 4-DOF or hybrid serial-parallel 5-DOF 4-coupled-Cartesian-manipulator.", + "texts": [ + " 10 ( R P PP U ) ( P P P U ) \u2217 Extension to family of coupled-Cartesian-manipulators to investigate in further research \u2217 Enables rotation control of angle \u03b8T Z around common-link L T longitudinal \u02c6 Z T axis and D) In Sm illustrate manipulators with intersecting-revolute-axes. Each figure illustrates a stand-alone manipulator, with its own unique features that are summarized in Table 2 . The progression of Figs. 1\u20136 illustrates the hierarchical composition of the 5-DOF 3 T 2 R fully parallel mechanism of Fig. 6 . Two serial Cartesian manipulators Fig. 1 form the two hybrid serial- parallel manipulators of Figs. 3 , 4 which together form the single fully parallel 4-DOF 2 T 2 R manipulator Fig. 5 . Adding an active moving-base (base-link) extends it to the 5-DOF 3 T 2 R fully parallel manipulator Fig. 6 . To make it easier to see the mechanisms, only one out of four possible vertical supports L w is shown in Figs. 1\u20136 A. In practice, these mechanisms would typically have four vertical supports and redundant prismatic P 2 joints to evenly balance loads as shown in Figs. 6 B, 8 . The mechanism in Fig. 1 D represents a conventional ( P P P RRR ) serial Cartesian manipulator with a 3-DOF spherical wrist", + " The 5-DOF hybrid serial-parallel 2-coupled-Cartesianmanipulator with parallel-revolute-axes of Figs. 3 , 4 have joint topology ( PPP U )(P PP U ) . Note that they have four revolute R joints (two R \u2019s per U ) compared to the non-parallel-revolute-axes configuration Fig. 2 , with a total of five revolute R joints. The manipulator in Fig. 4 is identical to the one in Fig. 3 except that it is rotated \u03b8A 2 Z = \u221290 \u25e6, \u03b8C 2 Z = \u221290 \u25e6 around the \u02c6 Z W axis. The link and coordinate symbols in Fig. 4 are labeled with subscript \u2018 2 \u2019 to distinguish them from the ones in Fig. 3 with subscript \u2018 1 \u2019. Fig. 5 fully parallel 4-DOF 4-coupled-Cartesian-manipulator . Common-link L T in Fig. 5 connects the two 2-coupledCartesian-manipulators from Figs. 3 , 4 . Together they form the 4-DOF fully parallel 4-coupled-Cartesian-manipulator of Fig. 5 with 2 T 2 R motion-type. The passive revolute R joint, along the common-link \u02c6 Z T axis, accommodates parasitic-twistangle \u03b8BD n Z between links L B 1 , L D 1 and L B 2 , L D 2 since the revolute R joint axes of the two manipulators from Figs. 3 , 4 are not geometrically parallel to each other. Coaxial revolute R 1 , R 2 , R 3 joints constrain the orientation of the common-link \u02c6 Z T axis. In practice R 1 , R 2 may be a matched pair of coaxial, separated, angular-contact bearings forming a single revolute R joint. Faint dotted lines in Fig. 5 A Ni Sc depict a possible coaxial revolute R 3 joint. However, it is not required in practice and over constrains the other coaxial revolute R 1 , R 2 joints. Four parallel-connected actuators, controlling the linear positions x A 1 A 1 , x C1 C1 , x A 2 A 2 , x C2 C2 of links L A 1 , L C 1 , L A 2 , L C 2 adjust the linear position x W T , y W T of the common-link L T and angular orientation of its \u02c6 Z T axis with 4-DOF and 2 T 2 R motion-type. An important distinction of the 4-coupled-Cartesian-manipulator is that all of the active prismatic P joints transmit axial forces directly to the common-link L T as opposed to some of the active prismatic P joints transmitting lateral forces to the common-link L T for the 2-coupled-Cartesian-manipulators in Figs. 2\u20134 . Short arrows identify active prismatic P joints in schematic In Sc of Fig. 5 B. Linear position z W A 2 = 0 is fixed for the 4-DOF manipulator in Fig. 5 , with joint topology (P P RRR )3( PP P RR ) . Notation \u20183( \u00b7 )\u2019 indicates three limbs ( \u00b7 ) connected in parallel. The extra R in the first limb ( P P RRR ) accommodates the parasitic-twist-angle \u03b8BD n Z defined below. The first limb ( P P RRR ) Please cite this article as: P. Wiktor, Coupled Cartesian manipulators, Mechanism and Machine Theory, https://doi.org/10. 1016/j.mechmachtheory.2020.103903 P. Wiktor / Mechanism and Machine Theory xxx (xxxx) xxx 11 has only two prismatic joints since it is fixed along the vertical \u02c6 Z W axis. The intersecting-revolute-axes version in Fig. 5 B, 5D has joint topology (P P UR )3( PP P U) , which may be represented as follow, suggesting the name \u2018Xactuator\u2019: Controlling the prismatic joint linear position z W A 2 in Fig. 5 B, 5D along vertical link L W adds position control z W T along \u02c6 Z W of the common-link L T for 5-DOF in a hybrid serial-parallel configuration, with joint topology ( P P P UR )3( PP P U) . In Fig. 5 , the 2-coupled-Cartesian-manipulator from Fig. 3 interleaves with the one from Fig. 4 so that link L B 2 from the second 2-coupled-Cartesian-manipulator is between links L B 1 , L D 1 from the first 2-coupled-Cartesian-manipulator. Similarly, link L D 1 from the first 2-coupled-Cartesian-manipulator is between links L B 2 , L D 2 from the second 2-coupled-Cartesianmanipulator. This allows coaxial bearings R 1 , R 2 to be spaced far apart from each other, along the common-link L T , as seen in Fig. 5 A, to support high moment loads with high moment stiffness. The load support link in solid model In Sm Fig. 5 D supports external loads applied to the common-link L T and the weight of the links and joints along the common-link L T . A redundant passive ( PPRP ) limb connects the load support link to the workspace link L W . Its revolute R joint allows the load support link to swivel with changes in orientation of the common-link L T . Fig. 6 fully parallel 5-DOF 4-coupled-Cartesian-manipulator with active base-link L Bs . The two parallel-connected manipula- tors In Sc , In Sm in Fig. 6 A, 6B control the 5-DOF position and orientation of common-link L T relative to base-link L Bs " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001597_j.cad.2019.102798-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001597_j.cad.2019.102798-Figure2-1.png", + "caption": "Figure 2: Angles Nomenclature. For any point p = pi, pprev refers to the point p(i\u22121)%n while pnext refers to the point p(i+1)%n. The point p\u2032prev is the reflection of pprev at p. The angle \u03b1p, which is 180\u25e6 \u2212 \u03b2p, is then formed by the triplet points of {pnext, p, p\u2032prev}, while \u03b2p is formed by {pprev , p, pnext}.", + "texts": [ + " There are three types of breakpoints: large angle breakpoints, curvature based breakpoints, and bounding box breakpoints. Large angle breakpoints are used to capture the main feature200 points of the 2D polygonal boundary, curvature-based breakpoints are used to place points at a constant interval around curved regions to form an approximation of the curved feature, and bounding box breakpoints are used to ensure that the deviation from the original boundary is kept within a reasonable range. For clarity, Figure 2) illustrates the convention used while referring to angles205 formed by three consecutive points. This nomenclature will be used henceforth in this paper. Large Angle Breakpoints Large angle breakpoints are characterized by any point p with \u03b1p > 30\u25e6. The point p is then used as a breakpoint to split the chain of points into smaller210 segments. Assuming that m number of breakpoints exist, represented by the index set b = {b1 . . . bm}, segments can be defined as S = {S1 . . . Sm}, where 9 Jo ur na re -p oo f Si = {pbi , pbi+1, " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure3-1.png", + "caption": "Figure 3 Constraint position", + "texts": [ + "25 (1) \ud835\udc47 = 2\ud835\udc53\ud835\udc41\ud835\udc45 (2) N: The pressing force of the friction pad on the brake disc, N; P1: the pressure of the hydraulic brake pipeline, MPa; S1: the area of the pad, cm2; d1: the piston diameter, cm; f: the friction coefficient, R: Take the radial dimension, cm. Substituting the known data into the above formula can get Nmax=18096 N. Adding load, the disc brake mainly acts on the brake disc through the friction block when the load is applied, and the load is applied to it as shown in the figure below (Figure 3). Defining constraints, apply full constraints on the center hole surface of the brake disc, apply X constraints on the inner and outer end surfaces of the disc, and apply Y and Z constraints on the friction surface of the brake disc and the friction lining. Its structure is shown in the figure below (Figure 4): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 The simulation calculation and analysis of the structural strength of the brake disc, and the results of stress solution and modal analysis deformation cloud diagram are shown in the following figure (Figure 5-6): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000946_icma.2019.8816196-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000946_icma.2019.8816196-Figure1-1.png", + "caption": "Fig. 1 Manipulator structure diagram", + "texts": [ + " Aiming at the problems that the model parameters of the manipulator system and the load range are varied, a fuzzy sliding mode control algorithm is proposed. Adding 25 fuzzy logic rules to the fuzzy system to eliminate the uncertainty in sliding mode controller, the method effectively eliminates the buffeting phenomenon. The simulation results show that the controller has a high control precision and strong robustness, which is sufficient to meet the requirement of positioning accuracy for the projectile in engineering. Fig. 1 is the structure of a transfer manipulator. Its mechanical system consists of three parts: base, swivel and translation. The swivel part is mainly composed of rotary drive shaft, drive gear and rotary body. The drive gear is connected to the permanent magnet synchronous motor through a rotary drive shaft and a speed reducer, it provide input torque for the rotary motion of the manipulator. Before establishing the mathematical model of the manipulator, the system should be simplified[4]. Fig. 2 is a 822978-1-7281-1699-0/19/$31" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003635_052061-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003635_052061-Figure7-1.png", + "caption": "Figure 7. Results FEM contact simulation (bottom outer ring to top inner ring)", + "texts": [ + " \u2219 \"# \u2219 $\u20d7# , $\u20d7# , % $\u20d7& , (1) Thereby, $\u20d7# , $\u20d7& are the contact radii towards the bearing center axis, \"# is the rotational speed of the inner ring and is the rolling diameter. The kinematic model neglects the displacement of the rolling elements geometric center due to their elastic deformation towards the center of the raceway curvature $\u20d7 '. Considering this effect requires an integration of the flattening effect due to the compression of the rolling elements by means of hertzian theory or alternatively by finite element contact simulation. Figure 7 (a) shows a comparison of the contact angles at the upper row (RW1) for a ring load distribution in negative direction of the blade root coordinates. Starting from the mounted contact angle at 47\u00b0 both angles deviate while arising due to different deformations of the outer ring (OR) and the inner ring (IR). Thereby, the contact angle is obtained by means of finite element simulation of the force vectors between the ring planes orthogonal towards the rotational bearing axis displayed in Figure 7 (b). There are various existing methods to calculate the cage speed either couple the cage to the rolling elements or to the shear effect of the lubricant. Since the contributed model currently neglects tribological losses methods coupling the cage to the rolling elements are considered. In case of modelling the cage coupled to the rolling elements the cage speed can either be determined by the leading rolling element, or by weighting the rolling element normal forces propagated for high speed applications by Wagner and Krinner et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001672_978-3-030-35699-6_49-Figure17-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001672_978-3-030-35699-6_49-Figure17-1.png", + "caption": "Fig. 17. 3D diagram of a system with multiple coils [3]", + "texts": [ + " Moreover, in case of a non-deformable contact, the non-magnetic object would have to be very resistant to the very important instantaneous acceleration during the shock. In a further work, the structure of our coil gun will be optimized dividing the coil in several smaller coils activated sequentially. For example, scenarios using 2, 3 or 4 coils (each one having a half, a third or a quarter of the total turns of the initial coil), triggered sequentially by software or using a position sensor instead of a single coil as shown in Fig. 17 will be evaluated as in [2,7]. This will lead to have successively a maximum current on each coil when the rod is optimally placed in the coil, leading to a increased projectile speed. Eighteenth International Middle East Power Systems Conference (MEPCON), pp. 506\u2013511, December 2016. https://doi.org/10.1109/MEPCON.2016.7836938 2. Bencheikh, Y., Ouazir, Y., Ibtiouen, R.: Analysis of capacitively driven electromagnetic coil guns. In: The XIX International Conference on Electrical Machines - ICEM 2010, pp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002957_022039-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002957_022039-Figure5-1.png", + "caption": "Figure 5. Robot movement based on user input", + "texts": [ + " JICETS 2019 Journal of Physics: Conference Series 1529 (2020) 022039 IOP Publishing doi:10.1088/1742-6596/1529/2/022039 The second module intended to expose students with the application of selection statement such as ifelse and switch-case in C language. Thus, the hands-on exercise involves the activation of the robot upon receiving commands from user\u2019s smartphone. To perform this, students need to use scanf function to read the input given by user and implement if-else selection statement in C programming in order to perform a decision-making process. For example, as depicted in Figure 5, if the input given was \u20181\u2019, then the robot shall move forward direction, if the robot received \u20182\u2019 or \u20183\u2019, it shall turn right and left respectively. By doing this, students will be able to observe the immediate effect of the if-else and switch structure in real life application. Module 3 emphasized on the use of repetition statement such as do-while, while, and for, to replicate certain robot action or movement pattern. The tasks include moving the robot in forward direction for two seconds, freeze for one second and repeat the action for three cycle recurrently as illustrated in Figure 6" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002624_icrom48714.2019.9071860-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002624_icrom48714.2019.9071860-Figure1-1.png", + "caption": "Fig. 1. Schematics of the proposed flexible manipulator", + "texts": [ + " Two learning based approaches based on Artificial Neural Networks (ANN) and Fuzzy Logic is used to find the inverse dynamics of the manipulator and performances of them compared to each other. II. MODELING In this paper a novel soft manipulator based on the magnetostriction phenomena is introduced, the proposed model consists of two slim beams connected to each other by means of a magnetic shielding plate. Each beam is made up of two layers, one layer of steel which is laminated with a layer of Metglas 2605sc as magnetostrictive material and wrapped by an electric coil (Fig. 1). Once an electric current flows in the coil, a magnetic field is generated inside it, which causes the beam to undergo a shape change and deflection in its 241 978-1-7281-6604-9/19/$31.00 \u00a92019 IEEE Authorized licensed use limited to: University of Edinburgh. Downloaded on June 14,2020 at 05:27:27 UTC from IEEE Xplore. Restrictions apply. transverse direction. The dynamic model of each beam is derived by means of nonlinear Euler-Bernoulli beam theory and the position and orientation of the end-effector is achieved in terms of the applied electric current", + " Substituting the approximate solution (7) into (5), multiplying both sides in the shape function and then integrating over the length of the beam, the following reduced order model in terms of the generalized coordinates is obtained: (8) where upper dots denote derivative with respect to time and the coefficients of the equation are given as: (9) It is worth mentioning that we assume that the acceleration of the origin of the coordinate system x2-z2 is negligible so that the assumptions of inertial coordinate system holds. This implies that we apply the current corresponding to the first link and once the equilibrium position is achieved (the origin of the x2-z2 coordinate system has no acceleration) the current corresponding to the second link is applied to achieve the desired end effector position. Afterwards the end-effector position is attained by means of following transformation matrix which is based on Fig. 1.a. (10) The second coordinate system is attached to the tip of the first link with x2 axis along the normal of the cross section of the tip of the first link, so the angle \u03b8 between two coordinate systems and also the orientation of the end-effector can be achieved as: (11) B. Inverse dynamics In order to use a manipulator in real world, studying the inverse kinematics is of significant importance; however, except for rare and few simple cases it is very challenging to find a closed-form solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003159_ijnsns-2018-0005-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003159_ijnsns-2018-0005-Figure2-1.png", + "caption": "Figure 2: (a): Kinematics of RIP system [17]. (b): Experimental setup of RIP system.", + "texts": [ + " Singla: ANFIS Based System Identification of UA Systems \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x\u0307 x\u0308 \u03b8\u0307 \u03b8\u0308 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 1 0 0 0 0 \u2212m M g 0 0 0 0 1 0 0 (M +m) Ml g 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x\u0307 x\u0308 \u03b8\u0307 \u03b8\u0308 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 1/M 0 \u22121/Ml \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6Fx (3) After including the parameter of the linear inverted pendulum, Equation (3) can be rewritten as: \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x\u0307 x\u0308 \u03b8\u0307 \u03b8\u0308 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 1 0 0 0 0 \u22120.98 0 0 0 0 1 0 0 35.93 0 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 x\u0307 x\u0308 \u03b8\u0307 \u03b8\u0308 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 + \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 1 0 \u22123.33 \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6Fx (4) Consider an RIP system consisting of the pendulum of length Lp connected/hinged to the end of a rotary arm of length Lr, as shown in Figure 2. The rotary arm is directly actuated by the direct-drive rotary servo motor of input voltage Vm while the pendulum is underactuated and is indirectly controlled by the rotary arm. Let, Jp and Jr correspond to the mass moment of inertia of the pendulum and the rotary arm about their center of mass, respectively. Here, \u03d5 symbolizes the angular position of the rotary arm and \u03b8 represents the pendulum angle which is zero when the pendulum is at the vertically upright position and is considered positive in counterclockwise direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000881_ecc.2019.8795797-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000881_ecc.2019.8795797-Figure4-1.png", + "caption": "Fig. 4. Phase 1 of the evasion: approaching Pi.", + "texts": [ + " If for any of the angles \u2220PiEPj > 2\u03c0/3, the evader can exploit it even if this means approaching one of the farther pursuers - in this case, there will still remain a neighbor distant enough to implement phase 2 (of Algorithm 1). (Note that this is the reason for requiring two pursuers to be distant.) Hence the equal angle distribution is the worstcase for the evader. Second, the pursuers that are not approached have equal speeds to the evader. Although it is intuitive, we will show that this is a conservative approximation after describing the phases of the pursuit. Phase 1) As shown on Fig. 4, angles and distances remain constant except the distance di to the pursuer Pi approached. Phase 2) In this phase, Pi and E both move perpendicularly to \u2212\u2212\u2192 EPi with different speeds. This causes them (in continous time, with instant information) to move along circles, whose radii are proportional to each other as the velocities: Ri R = vp ve (14) where Ri and R are the radii of the circles the pursuer and evader are moving on, respectively (as shown in Fig. 5). This R radius appears in the theorem" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000594_s1068798x19060133-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000594_s1068798x19060133-Figure1-1.png", + "caption": "Fig. 1. Basic structure of lathe drive.", + "texts": [ + " The fourth step is to reduce the parameters already found and all the forces and torques in the actual system to a single shaft, according to the rules of machine theory. The shaft chosen is either the motor shaft or the output shaft (the spindle). The fifth step in modeling the torsional system is to analyze the eigenfrequencies of the system and select those that are most significant. Usually, these are the 2\u20134 lowest eigenfrequencies. Where necessary, the diagram of the system is simplified. The following method is used to calculate the dynamic characteristics of a torsional system consisting of four shafts in the gearbox of a lathe (Fig. 1). The point moments of inertia I, kgf m2, of the cylindrical bodies are calculated from the formula ( )\u2212 = = \u00d7 \u22128 2 2 1 10 \u00a0 \u00a0 , k i i i i I g L D d 6 where is the density of the material (7.8 g/cm3); k is the number of cylindrical sections into which the part is divided; is the length of section i, cm; is the external diameter of section i, cm; and is the internal diameter of section i, cm. The moment of inertia of the electric motor is Iem = 0.157 kgf m2. The distributed moments of inertia of the cylindrical shaft Id, kgf m2/m (the moments of inertia per unit length) are calculated from the formula where is the shaft\u2019s external diameter, cm; and is its internal diameter, cm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure82.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure82.4-1.png", + "caption": "Fig. 82.4 SS GCI and AlSiC profile of von Mises stress (Model 1)", + "texts": [], + "surrounding_texts": [ + "In the coupled analysis, the thermal load was coupled with the structural load to find out the combined effect on brake disc models. The temperature induced at various time points was imported into static structural, and then structural loads were applied. The analysis was run for 36 s, i.e. the time taken by the vehicle to stop due to the application of the emergency brake. The output results of von Mises stress and total deformation developed in the model were recorded (Figs. 82.7, 82.8 and 82.9). Some important points that can be drawn from the analysis are: 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 979 \u2022 When compared to the Factor of Safety offered by the GCI models, the AlSiC models offer higher Factor of Safety. \u2022 For the same applied load, the AlSiC models have lower thermal stresses than the GCI models, as AlSiC material has greater thermal conductivity and heat dissipation capability. \u2022 The weight of AlSiC models is lesser when compared to the GCI models (GCI model having a weight of about 134 kg gets reduced to 54 kg in case of AlSiC material). 980 E. Madhusudhan Raju et al." + ] + }, + { + "image_filename": "designv11_80_0001956_vppc46532.2019.8952225-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001956_vppc46532.2019.8952225-Figure1-1.png", + "caption": "Fig. 1. Motor configurations. (a) Model 1. (b) Model 2. (c) Model 3. (d) Model 4. (e) Model 5. (f) Model 6.", + "texts": [ + " Section III compares the performance of these motors and Section IV focuses on the operation of the motors under prescribed driving cycles. Finally, some conclusions are given in Section V. II. MOTOR CONFIGURATIONS AND MATHEMATIC MODELS The conventional I-shape PM and novel cut-off rotor [24] structures for PMSM and SynRM are studied, where the winding layouts as well as the diameters and material of the stators and rotors are the same. The stack lengths may vary in order to obtain similar torque output. Six models, broken down into two categories, are illustrated in Fig. 1, where the motor parameters are presented in Table I. Note that Model 1 978-1-7281-1249-7/19/$31.00 \u00a92019 IEEE to 3 are based on the PMSM (specifically interior PMSM) and Models 4 to 6 are based on the SynRM. Among these models, Models 3, 5 and 6 are flux intensifying motors (FIMs) by adding flux barriers to q-axis from a PMSM (Model 3) or by adding PM in d-axis from the SynRMs (Models 5 and 6). In contrast, Models 1 and 2 are flux weakened motors (FWMs) where the detailed mathematical models will be described in next sub-section", + " Thus, the focal points of the torque hyperbola asymptotes and the maximum torque per ampere (MTPA) operations fall into the first quadrant. - For flux intensifying, with the added flux barrier on qaxis (Model 3), the major axis of the voltage limit ellipse for interior PMSM, i.e., purple ellipse, transforms from horizontal one [Fig. 3(a)] to vertical one [Fig. 3(b)]. In contrast, the voltage limit ellipses of SynRM, i.e., the green ellipses, move to the left with added PM on d-axis (Models 5 and 6). - As can be seen in Fig. 1, adding flux barrier on q-axis of PMSM leads to a decrease of Lq to be smaller than Ld. This, thus, becomes FIM. However, this may not make a large d-qaxis inductance difference or saliency (as shown in Fig. 3(b) where the purple ellipse is close to a circle) so that the reluctance torque and flux intensifying are limited. This will be explained in next section. III. MOTOR OPERATING CONDITIONS Fig. 4 presents the torque components (PM and reluctance torques) when applying peak current (120 A) to all the motor models", + " The PM torque is also much affected due to large leakage flux. The PM torque of Model 3 is higher than that of Model 2 and close to that for Model 1 and its reluctance torque is much lower than the other two cases. The reason is that the PM flux leakage of Model 3 decreases so that the PM torque increases. Increasing the saliency of Model 3 to increase reluctance torque is very difficult due to the limited free space near the rotor surface. In contrast, adding PM in d-axis to intensify the flux for SynRM is easy, as illustrated on Models 4 to 6 in Fig. 1. For Model 4 with only reluctance torque, adding PM and increasing PM width in d-axis would lead to a significant increase of PM torque and a decrease of reluctance torque, as shown in Models 5 and 6 in Fig. 4. Note that the decrease of reluctance torque is smaller than the increase of PM torque. Consequently, the total torque is much improved. However, the PM torque can become dominant and the motor would be more like PMSM if too much PM is used, in particular when the PM span (or width) is too big" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002571_icmre49073.2020.9065043-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002571_icmre49073.2020.9065043-Figure8-1.png", + "caption": "Figure 8. MSN navigation environment and models.", + "texts": [ + " The navigation capabilities of the MSNs is represented in the simulation by NI DaNI 1.0 MR CAD models. DaNI 1.0 is a four-wheeled MR featuring a reconfigurable FPGA board with real-time processors, sensors and other supporting hardware. The navigation environment considered is 8\u00d78m in dimension. In LabVIEW occupancy grid representation, each cell can be considered to be 0.5\u00d70.5m in dimension, making the environment a 16\u00d716 cells occupancy grid map. Each cell has a cost of \u2018100\u2019 representing an obstacle, storage rack or wall, and \u20181\u2019 representing free space. Figure 8 shows the MSN navigation environment model with two zones, zone 1 \u2013 ZSN1, ZSN3, MSN1 and zone 2 \u2013 ZSN5, ZSN7, MSN2. The ZSNs and MSNs default locations presented in Table II. The system simulation proceeds from the PCDE where pre-collected data of ZSNs and MSNs sensor libraries are processed and classified as detected data and events by the data processor. The detected data and events are displayed by the indicators, encoded and communicated via the network emulator from ZSNs to MSNs and MSNs to BSG" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000280_s10999-019-09455-z-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000280_s10999-019-09455-z-Figure5-1.png", + "caption": "Fig. 5 Meshing of star-ring and boundary conditions for type-C model at n \u00bc 0", + "texts": [ + " We estimate the tangential forces that act along the radial direction towards the centre of the ring by multiplying each elements of normal thrust force with a suitable coefficient of friction for metal to metal contact. We predict that these tangential forces resist deformation of the ring in the radial direction. In FEM modelling we apply these tangential forces in a vector format on the surface area between two bolts. The detail calculations are shown in the \u2018\u2018Appendix 2\u2019\u2019. Mesh generation, loading and boundary conditions are shown in Fig. 5a\u2013c. 4.2 Modelling procedure We consider the detail geometries of the star and ring and material properties of a specific commercially available ORBIT motor (Table 2) in FEM modelling. However, we consider that the star lobe and the ring lobe profiles are ideal neglecting the manufacturing tolerances. The fitting is also ideal. This means that at no load all the form-closed active contacts exist. But there is no contact force acting and no deformation or gap. In reality this condition is not achieved due to the manufacturing tolerances", + " For type B model those are illustrated in Fig. 3d\u2013f. We first estimate the amount of side thrust and the frictional forces on the ring plate, due to fixing bolt tightening forces, for type C FEM model. Detail calculations are shown in the \u2018\u2018Appendix 2\u2019\u2019. Figure 4 illustrates the directions of frictional forces and their distribution pattern, on one side of the ring. The loading and its pattern on the other side of the ring is exactly same. Mesh generation, assigned loading and boundary conditions at n \u00bc 0 are shown in Fig. 5a\u2013c. Finally we apply the load due to fluid pressure on the star i.e., the rotor, at the imaginary plane Li b and perform FEM computation. Figure 6a illustrates the deformations and gaps at transition contacts, estimated by FEM type C (final) model, at lobes 1, 4 and 5, for different angular position (for 1st power cycle i.e. n = 0 \u20138.57 ) of the output shaft (rotation in clockwise direction), at 11 MPa differential pressure. Figure 6b demonstrates the deformations and gaps at other active contacts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure1.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure1.1-1.png", + "caption": "Figure 1.1 The scheme of a cantilever beam.", + "texts": [ + " Thus, we will have, omitting for simplicity asterisk at x: \ud835\udc59\ud835\udc64(x) = \u2211 i giA2n\u2212i+1\ud835\udeff(x \u2212 a) + \u2211 j gjA2n\u2212j+1\ud835\udeff(x \u2212 b). (1.9.26) Thus, the boundary value problem (1.9.22, 1.9.21) is equivalent to the boundary value problem with nonhomogeneous differential equation in (1.9.26) and the homogeneous boundary conditions at the ends of the interval \u03a9. The most important is the case where \u03a9= [a, b]. In this case Aiw(a \u2212 0) = 0 Ajw(b + 0) = 0 (1.9.27) and the equivalent operator has a \u201cstandard\u201d form (1.9.26, 1.9.27). As an example, consider the problem of bending of a cantilever beam by force P0 and by moment M0 (Figure 1.1a) [269]. The equation of equilibrium of a beam we will write in the form: \ud835\udc59\ud835\udc66(x) = [\ud835\udc38\ud835\udc3c(x)y\ud835\udc3c\ud835\udc3c(x)]\ud835\udc3c\ud835\udc3c = 0 (a < x < l) (1.9.28) 1.9 General Approaches to Constructing Boundary Equations 77 where EI(x) \u2013 variable rigidity of the beam. The boundary conditions are: y(+0) = 0; yI(+0) = 0; \ud835\udc38\ud835\udc3c(x)y\ud835\udc3c\ud835\udc3c(x) = M0; [\ud835\udc38\ud835\udc3c(x)y\ud835\udc3c\ud835\udc3c(x)]I = P0; (x = l \u2212 0). (1.9.29) In this case p0(x)=EI(x), and p1(x)= p2(x)= 0. According to (1.9.6) when \u03a9 = [0, l], the boundary-value problem (1.9.28, 1.9.29) is equivalent to the boundary-value problem for differential equation: \ud835\udc59\ud835\udc66(x) = M0[\ud835\udeffI(x \u2212 l)] + P0\ud835\udeff(x \u2212 l), (1.9.30) with the boundary conditions y\ud835\udc3c\ud835\udc3c(x) = 0, [\ud835\udc38\ud835\udc3c(x)y\ud835\udc3c\ud835\udc3c(x)]I = 0, (x = l + 0). (1.9.31) The boundary forms entering in (1.9.29) form the full system 2n = 4 of boundary conditions. The inhomogeneous boundary-value problems in the form (1.9.30) are often used in applications. We note only that the correct consideration of the problem demands the correct statement of the boundary conditions in (1.9.29) and (1.9.31). In the case of inhomogeneous conditions at the left end (Figure 1.1b) y(+0) = y0, yI(+0) = y0 I , (1.9.32) and homogeneous conditions at the right end, we will have \ud835\udc59\ud835\udc66(x) = y0[\ud835\udc38\ud835\udc3c(x)\ud835\udeff\ud835\udc3c\ud835\udc3c(x \u2212 l)]I + yI 0\ud835\udc38\ud835\udc3c(x)\ud835\udeff \ud835\udc3c\ud835\udc3c(x \u2212 l) (1.9.33) or \ud835\udc59\ud835\udc66(x) = y0\ud835\udc38\ud835\udc3c(l)\ud835\udeff\ud835\udc3c\ud835\udc3c\ud835\udc3c(x \u2212 l) + [y0\ud835\udc38\ud835\udc3c I(l) + yI 0\ud835\udc38\ud835\udc3c(l)]\ud835\udeff \ud835\udc3c\ud835\udc3c(x \u2212 l) (1.9.34) under boundary conditions y(\u22120) = 0, yI(\u22120) = 0. (1.9.35) 78 1 M E T H O D S O F D Y N A M I C D E S I G N O F S T R U C T U R A L E L E M E N T S The similar construction of the boundary equations can also be made for multidimensional problems. Let us further proceed to Cauchy\u2019s problems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001087_ccdc.2019.8832450-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001087_ccdc.2019.8832450-Figure1-1.png", + "caption": "Fig 1. A single track model of CTCs", + "texts": [ + " sin[ arctan( )], 1, 1, 2, 2,yi i Ni fi i iF F E B i f r f r (1) where yiF is the lateral force of each axle of CTCs; i is the adhesion coefficient of each tire; fiE is curvature factor; iB is rigidity factor; i is slip angle; NiF is the vertical load of each axle and it can be obtained by the mechanical equilibrium equation. 2.2 A Single Track Model of CTCs The single track model is widely applied in the research of the stability investigation of CTCs. The traditional single track model only considers the yaw motion and lateral motion, neglecting the rolling and pitching. As a result, it has only three DOFs. The force condition of the single track is shown in Fig 1. In the model, m1 and m2 denote the masses of the towing car and the trailer while Iz,1 and Iz,2 denote their yaw moments of inertia respectively. vx is the speed of the CTCs. The front wheel steering angle is \u03b4w. \u03b1f,1 and \u03b1r,1 are the front and rear axle side-slip angles of the towing car, respectively. C\u03b1f,1 and C\u03b1r,1 are the front and rear axle cornering stiffness of the towing car. Fyf,1 and Fyr,1 denote the front and rear axle lateral forces of the towing car. Fyh,1 and Fyh,2 denote the forces acting on the hitch point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003911_2374068x.2020.1835016-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003911_2374068x.2020.1835016-Figure1-1.png", + "caption": "Figure 1. Different views of hip implant.", + "texts": [ + " So, this research work initially focused on fabricating PU replicas from the combined process of FDM, CVS and VC and then utilises them in the IC process for the fabrication of biomedical implant. During the RIC process, the effect of two input factors, namely drying time (primary coat) and layer numbers (thickness of mould) on the hardness and microstructure of casted implants has also been investigated. Further, mathematical model for hardness has been developed by using Buckingham\u2019s Pitheorem. Hip joint (refer Figure 1) was taken as benchmark component in this research. Creo 2.0 software was utilised for preparing the CAD model and then transformed into STL format. FDM technique utilised such STL file to prepare the ABS-based master pattern, followed by the smoothing process to improve their surface quality & to reduce the staircase effect. After that, an indirect rapid tooling technique named as VC process has been employed with an aim to create multiple PU replicas to reduce the cost of AM for batch production" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001629_012016-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001629_012016-Figure7-1.png", + "caption": "Figure 7. Illustration of stress distribution for orientations in (a) 0\u00b0 (b) 45\u00b0 and (c) 90\u00b0 direction.", + "texts": [ + " In the experimental test, 5 specimens were tested for each orientation, except the case in Figure 6(b), in which test 5 was observed unjustified. In all the considered cases, the results show the same trend. As normally expected the maximum tensile stress is achieved for loading along printing direction (0\u00b0), while the lowest maximum tensile stress is obtained for loading in transverse direction (90\u00b0). The corresponding strain values at maximum tensile stress are 0,04, 0,035 and 0,018 for 0\u00b0, 45\u00b0 and 90\u00b0 respectively. The results of the finite element analysis using ANSYS 17.0 workbench are given in Figure 7 (a) \u2013 (c) for 0\u00b0, 45\u00b0 and 90\u00b0 orientations respectively. In the simulation of parts produced by 3D printing, which are identical with those produced for experimental test, orthotropic material properties are considered. The corresponding material property for FEA modeling is taken from previous study conducted by EI-Gizawy et al. [16]. The boundary conditions for the FEA are set in similar way as the experimental setup, i.e. one end of the specimen is fixed and the load is applied at the other end" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002434_iccad46983.2019.9037969-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002434_iccad46983.2019.9037969-Figure1-1.png", + "caption": "Fig. 1. Posture of non-holonomic wheeled mobile robot", + "texts": [ + " MODELLING OF DIFFERENTIAL DRIVE ROBOT Modeling of a differential drive robot is a task of presenting the overall model of the two-wheeled mobile robot based on the kinematic and dynamic models in addition to the DC motor dynamics that must be taken into account. In fact, the control inputs of the robot dynamic model are the torques delivered by the two DC motors incorporated in the left and the right wheels. A. Kinematic Modelling The typical model of non-holonomic wheeled mobile robot [9] is shown in the Fig.1. The robot is operating with two wheels in addition to a castor one ensuring its stability. Two different coordinate frames are presented: A global reference frame in which the robot moves in. It is denoted as and a local frame which is attached to the robot and is denoted as . The origin of the local robot frame is the mid-point Or on the axe between the wheels. 978-1-7281-2292-2/19/$31.00 \u00a92019 IEEE Authorized licensed use limited to: University of Wollongong. Downloaded on May 30,2020 at 23:58:12 UTC from IEEE Xplore" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003955_ecce44975.2020.9235383-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003955_ecce44975.2020.9235383-Figure10-1.png", + "caption": "Fig. 10. Dimensioned stator electrode array", + "texts": [ + " Lastly, these voltages are sent through linear high voltage amplifiers. To convert from displacement sensor outputs, d1, d2, d3 to zo, \u03b1, and \u03b2, a geometrical relationship is defined as follows z0 = d1 + d2 + d3 3 \u2212 d0 \u03b1(t) = \u2212d1 + 2d2 \u2212 d3 3rl \u03b2(t) = \u221a 3 (d1 \u2212 d3) 3rl (26) where d0 is the equilibrium gap distance and rl represents the outward distance to each position sensor from the center of the disk. The sensors may be positioned between the spacing of the outer electrodes, allowing for maximum electrode area, as shown later in Fig. 10 To determine the validity of the above analysis, and more generally, the feasibility of macro-scale electrostatic bearings, an experimental system has been designed. The system consists of a 130 mm diameter aluminum disk rotor, enclosed by two stator electrode arrays that are shown in Fig. 10. The stator outer and inner radii are dimensioned to match that of the rotor, to allow for equal charge distribution. The experimental conditions of the bearing are given in Table I. The experimental apparatus developed is shown in Fig. 11. To create the conductive electrodes, the stators of Fig. 10 take advantage of inexpensive printed circuit manufacturing and the rotor is simply a repurposed aluminum computer hard disk drive. Three laser displacement sensors, placed above the top stator, between the electrodes are used to determine the position of the rotor. Initial rotor air-gap and orientation is set using three fine adjustment screws. The whole apparatus will be placed in a translucent vacuum chamber to negate the breakdown limitation and allow for easy viewing of the levitation. The control strategy will be implemented on an FPGA, whose analog output will be connected to eight highvoltage DC amplifiers" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000155_fie.2018.8659280-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000155_fie.2018.8659280-Figure3-1.png", + "caption": "FIGURE 3", + "texts": [ + " Safety concerns form an integral part of electronic design and test and were enforced throughout this workshop. Upon placement of the fan blade on the motor (M1), and closure of the slide switch (S1), the fan will spin and the lamp (L1) should turn on. The light helps protect the motor from getting the full voltage when the slide switch is closed. The participants removed the fan and noticed how the lamp gets dimmer when the motor does not have to spin the fan blade. The parallel circuit using the lamp and fan is shown in Figure 3. In this connection, the lamp does not change the current to the motor (M1). The participants removed the fan and noticed how the lamp does not change in brightness as the motor picks up speed. The lamp has its own path to the battery (B1). Project #2: Transistor circuits Project #2 engaged the participants in the assembly of basic transistor circuits and their applications. Figure 4 illustrates the set-up of the PNP collector circuit to demonstrate the effect of gain control (using the variable resistor, RV) on the operation of the transistor" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001552_s12555-019-0234-y-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001552_s12555-019-0234-y-Figure1-1.png", + "caption": "Fig. 1. Required output rolling motion (blue) and actual input motion (red) for laparoscopic suturing using (a) a straight wrist-less instrument and (b) a wristequipped instrument.", + "texts": [ + " Especially, distal axial rotation of the end-effector is very important for suturing with a curved needle. Because it is difficult to achieve distal rotation by using slender laparoscopic instruments, proximal rotation is generally transmitted to the distal tip with appropriate movement of the wrist joint [3]. As a result, unlike a joint-less straight instrument, a laparoscopic instrument equipped with a 2-DOF wrist joint permits the distal rotation required for suturing at various angles, as illustrated in Fig. 1. Among the various trials conducted to construct a small cable-driven wrist joint for a laparoscopic instrument, the one involving the use of an elastic backbone, such as a Manuscript received April 1, 2019; revised September 20, 2019; accepted September 25, 2019. Recommended by Editor Doo Yong Lee. Jungwook Suh is with the Department of Robot and Smart System Engineering, Kyungpook National University (KNU), 80 Daehakro, Bukgu, Daegu 41566, South Korea (e-mail: jwsuh@knu.ac.kr). c\u20ddICROS, KIEE and Springer 2019 spring or a tube, proved to be an easy approach to implement 2-DOF motion [4\u20137]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000520_022075-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000520_022075-Figure1-1.png", + "caption": "Figure 1. Schematic drawing of electron beam additive manufacturing.", + "texts": [ + " In particular, Electron Beam Melting, which uses electronic emitters, so-called electronic guns, as energy sources for melting in the high-vacuum medium. In this paper the creation of metallic structures made of 321 stainless steel and Ti-6Al-4V in the process of electron-beam additive technology is investigated. Influence of printing parameters on macrostructure, formation of defects on the surface of polymetallic structures. The samples under study were obtained using the Laboratory Experimental Installation for additive electron-beam production of metal wire products as shown in figure 1. The device reads data from a file containing a three-dimensional digital model and applies sequential layers of material wire. The contours of the model layers are drawn by an electron beam that melts the wire at the point of contact. Melting is performed in vacuum working chambers, which allows working with materials sensitive to oxidation. Eight layers were deposited. Layers were approximately 1 mm tall [11]. metallography, samples were cut using electro-discharge machining and were mechanically ground and then polished and pickled in HOOC-COOH solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003614_j.procir.2020.04.092-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003614_j.procir.2020.04.092-Figure11-1.png", + "caption": "Fig. 11. The intersection part (\u03a6\ud835\udc56\ud835\udc56) is bounded by { 180 \u2212 \ud835\udc65\ud835\udc65, \ud835\udc65\ud835\udc65 ,120 \u2212 \ud835\udc66\ud835\udc66 , \ud835\udc66\ud835\udc66 ,20 \u2212 \ud835\udc67\ud835\udc67 , \ud835\udc67\ud835\udc67}, where \ud835\udc67\ud835\udc67 = 20, \ud835\udc67\ud835\udc67 = 0, \ud835\udc66\ud835\udc66 = 120, \ud835\udc66\ud835\udc66 = 0, \ud835\udc65\ud835\udc65 = 180, \ud835\udc65\ud835\udc65 = 0.", + "texts": [], + "surrounding_texts": [ + "In this section, the proposed method is validated by a remanufacturing case study. There is a used part (Fig. 7a) required to remanufactured to a final part (Fig. 7b). The CAD models of the used part and final parts are modelled by discrete level set functions as \u03a6\ud835\udc62\ud835\udc62 and \u03a6\ud835\udc53\ud835\udc53 , separately. Both of them are built on a design domain D (150*150*100) with grid size \u2206\ud835\udc65\ud835\udc65 (0.5 mm), as shown in Fig. 8. The next step is the maximization of the intersection volume, so the intersection part extraction algorithm is implemented for the used part. As Fig. 9 shown, the used part is transformed from \u03a6\ud835\udc62\ud835\udc62 to \u03a6\u0303\ud835\udc62\ud835\udc62 by the optimal transformation matrix obtained by the proposed algorithm and then the intersection part \u03a6\ud835\udc56\ud835\udc56 is obtained. The result of the optimal transformation is shown in Table 3. After obtaining the intersection part, the next step is modifying the intersection part from \u03a6\ud835\udc56\ud835\udc56 to \u03a6\u0303\ud835\udc56\ud835\udc56 by respecting different AM processes. In the DED process, the material deposition nozzles have collisions with the part. The intersection part is modified from \u03a6\ud835\udc56\ud835\udc56 to \u03a6\u0303\ud835\udc56\ud835\udc56 by the implementation of Eq. (9)-(11). The result is shown in Fig. 10. In terms of the PBF process, the intersection part is modified automatically with respect powder recoater collision problem. The result is shown in Fig. 12, and the level-set function of the optimized intersection part is \u03a6\u0303\ud835\udc56\ud835\udc56 = min{ \u03a6\ud835\udc56\ud835\udc56, 180 \u2212 \ud835\udc65\ud835\udc65, \ud835\udc65\ud835\udc65 ,120 \u2212 \ud835\udc66\ud835\udc66 , \ud835\udc66\ud835\udc66 ,20 \u2212 \ud835\udc67\ud835\udc67 , \ud835\udc67\ud835\udc67}. The last step is individual features extraction. As shown in Fig. 12, a level set represented subtractive manufacturing volume is \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34. \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 is converted to \ud835\udc35\ud835\udc35\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 by the CSG to Brep conversion algorithm. Then, individual SFs (SF1 and SF2) are extracted by the machining feature recognition algorithm. The feature recognition step from additive manufacturing volume to individual features is represented in Fig. 13. The modified additive feature volume (\u03a6\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305) is divided into AFs and the residual subtractive feature volume (\u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 ). Similarly, the implicit form of subtractive feature volume is converted to B-rep. SF3-SF6 are then recognized." + ] + }, + { + "image_filename": "designv11_80_0001359_et.2019.8878654-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001359_et.2019.8878654-Figure2-1.png", + "caption": "Fig. 2 SRM linear inductance profile of a one phase", + "texts": [ + " The approach proposed in [3] and [6] for the synthesis of a linear model of a 6/4 SRM is used for the creation of the proposed 12/8 SRM specific model. The SRMs\u2019 main structures cause strong nonlinear magnetic characteristics, complicating its analysis and control. Frequently, the analysis takes into account that the motor remains magnetically unsaturated during operation. Then the phase inductances change linearly from the rotor position in the absence of magnetic saturation. In this case, a mathematical model of the SRM can be developed based on the equivalent circuit shown in Fig. 2, where R is the active winding resistance of the motor, and L is a phase inductance. The influence of the mutual inductance between the phases is neglected. Fig.1 SRM Single-phase equivalent circuit Since we stand in linear regime, when magnetic saturation is neglected, the relationship from flux \u03a6 to current is given by [1], ( ) iL \u22c5= \u03b8\u03a6 , (1) where L is the inductance dependent on the rotor position and phase current i. The electric equation, used to represent an instantaneous voltage across the terminals of a phase, is related to the flux \u03a6 linked in the winding as follows dt diRV \u03a6+= , (2) where V is the terminal voltage", + " For one phase, the co-energy is given by 2 ( ) 2cW i L \u03b8= , (5) The torque generated by a switched reluctance motor T can be approximated by: 2 2 cdW i dLT d d\u03b8 \u03b8 = = (6) This dependence shows that only in the absence of saturation does the inductance depend only on the rotor position angle . Under the simplified assumption of magnetic linearity, the total instantaneous torque equation becomes: 2 2total phase T i dL d\u03b8= (7) The following motion equations are used for receiving the SRM motor speed: \u03c9\u03c9 FTT dt dJ load \u2212\u2212= (8) dtd\u03b8\u03c9 = , (9) where J is the total inertia referred to the motor shaft, T is the motor torque, F is the friction coefficient, is the angular motor speed, loadT - load torque applied to the motor shaft. Fig. 2 shows the linear inductance profile ( )\u03b8L [1]. In order for each phase to be moved to the previous one, a multiphase angular profile is created. The linear inductance profile ( )\u03b8L , with each phase inductance displaced by an angle 1, is given by ( )[ ] 2)(21 rsrN \u03b2\u03b2\u03c0\u03b8 +\u2212= , (10) where rN is the number of rotor poles, s\u03b2 is the stator arc angle and r\u03b2 is the rotor arc angle. The next angels are defined as follows: s\u03b2\u03b8\u03b8 += 12 (11) )(23 sr \u03b2\u03b2\u03b8\u03b8 \u2212+= (12) 4 3 s\u03b8 \u03b8 \u03b2= + (13) rN\u03c0\u03b8\u03b8\u03b8 2145 =+= (14) In addition, to define the linear inductance profile, it is necessary to select a power transistor circuit to help simulate SRM performances" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002337_eiconrus49466.2020.9039215-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002337_eiconrus49466.2020.9039215-Figure4-1.png", + "caption": "Fig. 4 For integration over curved and rectilinear segments", + "texts": [ + " The limits of the external integral in (18) are taken by half of the element, since in (12) the integration region by the point Q is a union of half k -th and 1k + -th elements. Thus, calculating the integral in (18) by region kl\u2032\u2032\u0394 , then by calculating the integral over the region 1kl +\u2032\u0394 (and this region may be an arc of a circle), and summing results, we obtain the value of the integral over the union of the regions. The integrals in (18) are calculated numerically. Integration over curved and rectilinear segments. Fig. 4 shows the case when one of the elements over which the integration is carried out is rectilinear, and the second is an arc of a circle. Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM. Downloaded on June 14,2020 at 17:56:06 UTC from IEEE Xplore. Restrictions apply. 650 Consider the integrals: ( ) 1 2 4 2 , k i PQ P PQ pQ P Q P Q pQl l r n r r n dl dl r +\u2032\u0394 \u0394 \u2212 \u03c4 . (19) We divide the segment 1kl +\u2032\u0394 into km segments. Figure 4 shows the first such segment 1x\u0394 . Then the external integral in (19) can be replaced by the sum: ( ) 2 4 1 2 ,k s s s si m PQ P PQ pQ P Q P s s pQl r n r r n dl x r= \u0394 \u2212 \u03c4 \u0394 . (20) Integration by il\u0394 , as in the previous example, is replaced by integration by angle: P i Pdl d= \u03c1 \u03b1 . Vectors sQr can be defined like this: ( ) ( ) ( ) ( )1 1cos sin s s s s sQ Q Q k Q Q kr r e Q r e Q+ \u03c1 + \u03b1= \u03b1 \u2212 \u03b1 + \u03b1 \u2212 \u03b1 ; (21) ( )2 2 2 2 cos s s s sPQ Q i Q i Q Pr r r= + \u2212 \u2212\u03c1 \u03c1 \u03b1 \u03b1 . 1 1 0 1 1 0 1 2 2 2 1 1 2 2 2 1 2 sin ; arccos , " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001569_ecce.2019.8912739-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001569_ecce.2019.8912739-Figure7-1.png", + "caption": "Fig. 7 Equivalent ring model of the ASRM", + "texts": [ + " Equations of total maximum real stresses within each beam are summarized in Table I. In the ERM, the concept of transforming the active part of the rotor geometry to an equivalent ring with an equivalent mass density \u03c1eq is applied in order to find stresses in the air gap iron bridges [6]. The stress \u03c3eq inside the equivalent ring at a specific rotating speed is determined by (2). eqeqeq R \u03c1\u03c3 \u00d7\u03a9\u00d7= 22 (2) Req is the radius of the equivalent ring of thickness equal to ea, the thickness of the air gap bridge (Fig.7). This assumption explains the fact that the ERM is not directly applicable to the central iron bridge [7]. SCF are calculated using the ERM-SCF method proposed in [6] and applicable to round shapes flux barriers notches. Fig.8 defines the main geometrical parameters used in the ERM calculation. The analytical equations of the total stress within the air gap bridge are summarized in Table II. To sum up, at this stage, there are two corrected analytic models that allow the determination of high speed mechanical stresses on the ASRM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003614_j.procir.2020.04.092-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003614_j.procir.2020.04.092-Figure13-1.png", + "caption": "Fig. 13. The feature recognition process from additive manufacturing volume to individual features", + "texts": [ + " 12, and the level-set function of the optimized intersection part is \u03a6\u0303\ud835\udc56\ud835\udc56 = min{ \u03a6\ud835\udc56\ud835\udc56, 180 \u2212 \ud835\udc65\ud835\udc65, \ud835\udc65\ud835\udc65 ,120 \u2212 \ud835\udc66\ud835\udc66 , \ud835\udc66\ud835\udc66 ,20 \u2212 \ud835\udc67\ud835\udc67 , \ud835\udc67\ud835\udc67}. The last step is individual features extraction. As shown in Fig. 12, a level set represented subtractive manufacturing volume is \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34. \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 is converted to \ud835\udc35\ud835\udc35\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 by the CSG to Brep conversion algorithm. Then, individual SFs (SF1 and SF2) are extracted by the machining feature recognition algorithm. The feature recognition step from additive manufacturing volume to individual features is represented in Fig. 13. The modified additive feature volume (\u03a6\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305) is divided into AFs and the residual subtractive feature volume (\u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 ). Similarly, the implicit form of subtractive feature volume is converted to B-rep. SF3-SF6 are then recognized. Remanufacturing of worn products has considered as a promising strategy to reduce environmental impacts and financial expenditure by extending their production life cycle. By combining AM and SM, HM provides an efficient remanufacturing strategy, which allows the manufacture of a new part directly from a used part without the recycling process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure48.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure48.3-1.png", + "caption": "Fig. 48.3 Schematic of the 3D model of the microwave drilling", + "texts": [ + " But, it is very difficult to have a clear understanding of the process mechanism of microwave drilling due to the limitation in experimental assessment of distribution of electric field inside the cavity, temperature, and thermal stress along the surface of the workpiece, etc. So, the software tool like COMSOL Multiphysics can help in having a better understanding of the process mechanism. 48 Analysis on Thermal Characteristics of Micro-Drilled Glass \u2026 559 A 3D model of the experimental setup of microwave drilling as shown in Fig. 48.3 was developed using COMSOL Multiphysics 5.2 software to analyze the process and the behavior of workpiece while machining. The dimensions of the above model are shown in Table 48.1. 560 G. Kumar and A. K. Sharma Physics-controlled meshing is used to mesh the 3D model of the microwave drilling. Figure 48.4 shows the meshed 3D model. Fine meshing has been used to mesh the cavity (Fig. 48.4a), and extremely fine meshing has been used to mesh the concentrator and workpiece (Fig. 48.4b). The number of vertex elements, edge elements, triangular elements, and tetrahedral elements is 40, 557, 4504, and 45,401, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003739_s12239-020-0106-8-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003739_s12239-020-0106-8-Figure3-1.png", + "caption": "Figure 3. Schematic diagram about mechanical analysis of wheel external structure under driving condition.", + "texts": [ + " The following will carry out mechanical modeling for driving and braking conditions. 3.1. Mechanical Model of Modular Deformable Tire Under Driving Condition Under the driving condition, the driving torque Md is applied to the geometric center o of the tire, at which time the axle drives the wheel to roll forward. Under the action of car body gravity, the load acting on tire is H. Due to the joint action of the load H and the tire's own gravity G, the tire and the ground produce the friction fd as shown in Figure 3, and this friction is the driving force provided by the ground to the tire. Assuming that the acceleration and angular acceleration of the tire under driving state are a and \u03b1 respectively, then the mechanical equation of the wheel can be constructed: (1) where T is the supporting force of the ground to the tire; m is the quality of the tire; m is the mass of the tire; Fx is the horizontal force acting on the tire by the axle under the driving condition; J is the inertia moment of the tire; R is the moment arm of driving force; Mf is the rolling resistance moment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001623_ecce.2019.8912961-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001623_ecce.2019.8912961-Figure5-1.png", + "caption": "Fig. 5. Schematic of a cross-section of IM2", + "texts": [ + ", In denotes the identity matrix of order n. III. INDUCTION MACHINE MODEL INCLUDING WINDING DISTRIBUTION HARMONICS As the considerations in this paper rely on the model proposed in [5], we provide a brief summary of the underlying modelling hypotheses in this section. Stator windings are located in Ns slots and modelled as a set of ms electrical circuits. Such a circuit does not necessarily represent a complete winding as it might only consist of a subset of the coils belonging to this winding (see fig. 5 and 6). The rotor features Nr slots and mr electrical circuits. For squirrel cage machines, mr = Nr and each rotor circuit corresponds to a coil having one turn and a pitch equal to the rotor slot pitch. We presume that all stator circuits (resp. rotor circuits) have the same resistance Rs (resp. Rr) and the same leakage inductance L\u03c3s (resp. L\u03c3r). 978-1-7281-0395-2/19/$31.00 \u00a92019 IEEE 5293 As in [5], only space harmonics resulting from the conductor distribution are taken into consideration. Thus, we assume a linear magnetic behaviour and a constant air-gap length" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000276_ecai.2018.8679000-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000276_ecai.2018.8679000-Figure1-1.png", + "caption": "Fig. 1. Robot\u2019s position in both the global (XG-YG) and robot\u2019s local reference frame (XR-YR)", + "texts": [ + " Section II describes the kinematic model of the robot used in this paper while section III provides the details on the proposed new approach. Section IV discusses the results obtained by implementing this algorithm. II. KINEMATIC MODEL OF THE ROBOT Numerous mobile robots employ the differential drive mechanism to move from one place to another. To demonstrate the performance of the proposed algorithm, the differential drive robot which is capable of varying trajectories by merely varying the velocities of the two wheels is employed. Fig. 1 below shows the position of a robot in both the global and robot\u2019s local reference frame [2]. It is located at an arbitrary position (x,y) in a 2-dimensional space. It is moving in a direction that makes the angle \u03b8Ort with the XG axis of the global reference frame denoted by XG-YG. \u03b8Ort is the orientation angle of the robot. The robot\u2019s pose P based on the global frame can be described as in (1). The orthogonal rotation matrix R that maps motion along the global frame to that of the robot\u2019s frame is expressed as in (2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure21.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure21.1-1.png", + "caption": "Fig. 21.1 Weld pad model", + "texts": [ + " Numerical simulations were performed using commercially available SYSWELD software. The optimized meshing is necessary to resemble the exact experimental result in the numerical model. Hence, optimization of mesh size, with the right choice of mesh dimensions is required for improving the accuracy of the simulation. In this work, around 30 mm near to the weld line, fine mesh to a size of 1 mm 0.8 mm (quadrilateral) was chosen, since the effect of the temperature gradient is more near the weld region and heat affected zone as shown in Fig. 21.1. Goldak double ellipsoidal heat source model was used for defining a heat source [19]. The thermomechanical analysis is done by sequential coupling of thermal and mechanical analysis. Here the thermal and phase transformation strains are coupled 21 Numerical Simulation and Experimental Validation \u2026 247 with non-linear mechanical analysis to understand the evolution of stress and distortion in the weld joint. It was assumed that there is no influence of any structural changes on the thermal analysis part" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000289_cyber.2018.8688118-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000289_cyber.2018.8688118-Figure6-1.png", + "caption": "Fig. 6. The structure of the OMR", + "texts": [ + " With the purpose of features extraction of raw sEMG, an envelop of the raw sEMG signals from all Nc channels is extracted under the help of low pass filtering as well as moving average techniques. According to [22], moving average process bases on A(nt) = 1 Sw Sw\u22121\u2211 k=0 EMG(nt\u2212k) (1) where A(nt) is the amplitude of enveloped sEMG signal, Sw is the window size, EMG(nt\u2212k) is the sEMG sinal amplitude at simple time k and t is current time. By summing the absolute value of enveloped sEMG amplitude from each channel together, an index of coordinated muscle activation level can be calculated, defined as follows: P (k) = Nc\u2211 i=0 | Ai(nt) | (2) As shown in Fig.6, robot coordinate system is defined as xoy, the ith omnidirectional wheel coordinate system is defined asxioiyi. Define \u03b1ri = 90\u25e6 which is the angle between ith the wheel and its rollers, and \u03b1r1 = \u03b1r2 = \u03b1r3. According to Fig.6, the following equation can be obtained\u23a1 \u23a3vwix vwiy \u23a4 \u23a6 = \u23a1 \u23a3 0 sin\u03b1ri rw cos\u03b1ri \u23a4 \u23a6 \u23a1 \u23a3\u03c9w i vri \u23a4 \u23a6 = Ri 1 \u23a1 \u23a3\u03c9w i vri \u23a4 \u23a6 (3) where vwix and vwiy are the velocity of the ith wheel in xi axis and yi axis in xioiyi coordinate system respectively. \u03c9w i is the angle velocity of the ith wheel. vri is the velocity of the roller of the ith wheel. rw is the radius of the omnidirectional wheel. According to [23], mapping the ith omnidirectional wheel\u2019s velocity to xoy coordinate system, then\u23a1 \u23a3vix viy \u23a4 \u23a6 = \u23a1 \u23a3cos \u03b7i \u2212 sin \u03b7i sin \u03b7i cos \u03b7i \u23a4 \u23a6 \u23a1 \u23a3vwix vwiy \u23a4 \u23a6 = Ri 2R i 1 \u23a1 \u23a3\u03c9w i vri \u23a4 \u23a6 (4) vix and viy mean the velocity of the OMR in x axis and y axis in xoy coordinate system, which offered by the ith wheel", + " Then the inverse kinematics equations of the ith wheel of the OMR is \u23a1 \u23a3\u03c9i vir \u23a4 \u23a6 = [Ri 1] \u22121[Ri 2] \u22121Ri 3 \u23a1 \u23a3vxvy \u03c9 \u23a4 \u23a6 (7) define \u03b3ri = \u03b7i \u2212 \u03b1ri, then Ri = 1 \u2212rw sin\u03b1ri\u23a1 \u23a3cos(\u03b3ri) sin(\u03b3ri) \u2212lwiy cos(\u03b3ri) + lwix sin(\u03b3ri) r cos \u03b7i \u2212r sin \u03b7i lwiy \u00b7 r cos \u03b7i \u2212 lwix \u00b7 r sin \u03b7i \u23a4 \u23a6 (8) Define K1 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 cos(\u03b3r1) sin\u03b1r1 cos(\u03b3r2) sin\u03b1r2 cos(\u03b3r3) sin\u03b1r3 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 K2 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 sin(\u03b3r1) sin\u03b1r1 sin(\u03b3r2) sin\u03b1r2 sin(\u03b3r3) sin\u03b1r3 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 K3 = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 l1 sin(\u03b3r1 \u2212 \u03b21) sin\u03b1r1 l2 sin(\u03b3r2 \u2212 \u03b22) sin\u03b1r2 l3 sin(\u03b3r3 \u2212 \u03b23) sin\u03b1r3 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 (9) Combining Eqs.7 and i = 1, 2, 3, then the inverse kinematics model of the OMR can be obtained.\u23a1 \u23a2\u23a2\u23a2\u23a3 \u03c91 \u03c92 \u03c93 \u23a4 \u23a5\u23a5\u23a5\u23a6 = 1 \u2212rw [ K1 K2 K3 ] \u23a1 \u23a2\u23a2\u23a2\u23a3 vx vy \u03c9 \u23a4 \u23a5\u23a5\u23a5\u23a6 (10) The Jacobian matrix is R = 1 \u2212rw [ K1 K2 K3 ] (11) According to Fig.6 and the OMR used for the following experiment which is built by three omnidirectional wheels, \u03b1ri, \u03b2i, \u03b7i, r w and lw = lwi (i = 1, 2, 3) have been already fixed. The actual values are shown by table 1. kinematics model of the OMR is\u23a1 \u23a2\u23a2\u23a2\u23a3 \u03c91 \u03c92 \u03c93 \u23a4 \u23a5\u23a5\u23a5\u23a6 = \u23a1 \u23a2\u23a2\u23a2\u23a3 \u22120.0170 0.00987 2.3323 0 \u22120.0197 2.3323 0.0171 0.00987 2.3323 \u23a4 \u23a5\u23a5\u23a5\u23a6 \u23a1 \u23a2\u23a2\u23a2\u23a3 vx vy \u03c9 \u23a4 \u23a5\u23a5\u23a5\u23a6 (12) From Eqs.12, rank(R) = 3, which proves that the OMR can achieve omnidirectional movement and there is no nonholonomic constraints. Potential field method contains two artificial virtual fields, attractive potential field caused by goal and repulse potential field caused by obstacles" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure9.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure9.5-1.png", + "caption": "Fig. 9.5 A 3D finite element model DCR", + "texts": [ + " In this study, 56,250 elements and 58,420 nodes with 15 elements in thickness direction and 250 elements are used in the width direction. To save the time, a biased mesh is generated in the deformation region using double biasing with a bias ratio of 3. A roller wheel with 75 mm in diameter, 5 mm in thickness, 2.5 mm as nose radius is modelled as discrete rigid. The wheel is provided to have a uniform rotational speed of 10RPM. The simulations were conducted for penetration depths of 0.5, 1.0, 1.5 and 2.0 mm to deform the material longitudinally along the length of the specimen as shown in Fig. 9.5. Coulomb friction was given to the contact surfaces and it is important to note that adequate friction is required to contact between the sheet and deep roller. Frictional forces at the interface of the roller and the sheet are responsible for pulling the sheet through the roller. A residual stress gradient has been observed at the surface of the deep-rolled sheet. Boundary conditions are divided into two steps: an initial step and the analysis step. In the initial step, the sheet was given an initial velocity of 10 mm/s in 9 Experimental and Numerical Assessment of Residual Stresses \u2026 101 \u2018x\u2019-direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003043_jsen.2020.3007503-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003043_jsen.2020.3007503-Figure4-1.png", + "caption": "Fig. 4. Visual Occlusion situation during surgery.", + "texts": [ + " The 1f , 2f , 3f , 4f can be obtained by following equations: 2 2 1 1 2 1 1 L sin( ) 2 sin ( ) L sin( ) 2 sin ( ) 2 4 2 4 N N i i i i i i f d h d h = = = \u2212 \u2212 + \u2212 \u2212 \u2212 (11) 1 21 2 1 1 i 1 i i 1 1 ( ' 'sin( ) ( ' 0.5G )cos( ) 'cos( ) ) 2 N i i i i i x i y i i f M M L F F T \u2212 + + + + + = = \u2212 \u2212 \u2212 + + (12) 3 sin( ) ( ' 0.5G )cos( )n n ex n ey n nf M M L F F = \u2212 \u2212 \u2212 + (13) 4 ( )f b/N= \u2212bQ B (14) where bQ is the position information of the b-th joint measured by stereo vision system. During the surgery, only part of markers can be measured in most cases. As shown in Figure 4, the flexible manipulator will be blocked by other instruments or body organs. In this part, the experimental setup will be introduced firstly. Then, the simulation and experimental results for a serpentine flexible manipulator without external payload will be shown. After that, we will show the simulation and experimental results with external payload. The worst condition is considered (only one joint\u2019s position can be measured by stereo vision system) to show the performance of proposed shape sensing method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003614_j.procir.2020.04.092-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003614_j.procir.2020.04.092-Figure12-1.png", + "caption": "Fig. 12. The feature recognition process from subtractive manufacturing volume to individual features", + "texts": [ + " The result of the optimal transformation is shown in Table 3. After obtaining the intersection part, the next step is modifying the intersection part from \u03a6\ud835\udc56\ud835\udc56 to \u03a6\u0303\ud835\udc56\ud835\udc56 by respecting different AM processes. In the DED process, the material deposition nozzles have collisions with the part. The intersection part is modified from \u03a6\ud835\udc56\ud835\udc56 to \u03a6\u0303\ud835\udc56\ud835\udc56 by the implementation of Eq. (9)-(11). The result is shown in Fig. 10. In terms of the PBF process, the intersection part is modified automatically with respect powder recoater collision problem. The result is shown in Fig. 12, and the level-set function of the optimized intersection part is \u03a6\u0303\ud835\udc56\ud835\udc56 = min{ \u03a6\ud835\udc56\ud835\udc56, 180 \u2212 \ud835\udc65\ud835\udc65, \ud835\udc65\ud835\udc65 ,120 \u2212 \ud835\udc66\ud835\udc66 , \ud835\udc66\ud835\udc66 ,20 \u2212 \ud835\udc67\ud835\udc67 , \ud835\udc67\ud835\udc67}. The last step is individual features extraction. As shown in Fig. 12, a level set represented subtractive manufacturing volume is \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34. \u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 is converted to \ud835\udc35\ud835\udc35\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34 by the CSG to Brep conversion algorithm. Then, individual SFs (SF1 and SF2) are extracted by the machining feature recognition algorithm. The feature recognition step from additive manufacturing volume to individual features is represented in Fig. 13. The modified additive feature volume (\u03a6\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 \u0305) is divided into AFs and the residual subtractive feature volume (\u03a6\ud835\udc46\ud835\udc46\ud835\udc34\ud835\udc34\ud835\udc34\ud835\udc34\u0305\u0305 \u0305\u0305 \u0305\u0305 ). Similarly, the implicit form of subtractive feature volume is converted to B-rep" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001231_demped.2019.8864909-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001231_demped.2019.8864909-Figure10-1.png", + "caption": "Fig. 10: Picture of the prototype machine with full assembly.", + "texts": [ + " Figure 8 shows the differential reactive power as a function of time, in case of 10 % (blue line) and 20 % (red line) of shorted turns in a phase of three-phase system prime. Figure 9 shows the differential reactive power as a function of the percentage of shorted turns. Results confirm the model presented in previous section and that this diagnostic index can be an effective and robust tool to detect stator faults, even with a more realistic percentage of shorted turns (1-5%). Experiments were performed, with a prototype machine with 48 slots, Figure 10. Tests were performed supplying the three-phase systems with a power converter forcing perfectly sinusoidal currents. Faulty conditions were tested, so far, shorting the three lines of a three-phase system, while the other one is supplied with the power converter. In healthy and faulty conditions the machine under test is mechanically coupled to a controlled load spinning at 36 rpm, Figure 11. The phase currents in the shorted three-phase system were measured and they are reported in Figure 12. The proposed procedure can be successfully applied to this case of extreme fault" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000013_ccee.2018.8634528-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000013_ccee.2018.8634528-Figure2-1.png", + "caption": "Fig. 2: Mesh discretization on the SRM.", + "texts": [ + " In the non-linearity case the ferromagnetic material should specify by B-H curve as shown in Figure 1 where effects of magnetic saturation are taking in to the count in order to obtain static characteristics based on the use of finite element modeling. Electromagnetic modeling by the MEF requires the definition of the magnetic permeability for each zone 978-1-7281-0112-5/18/$31.00 \u00a92018 IEEE according to the properties of the material or the space. In Figure 1 the magnetization curve of the material are presented for the yoke of the rotor and the stator. From the above, we find that in this section we need to define the geometry of the SRM in the ANSYS software tools in order to design the SRM section based on the data provided. In Figure 2, we have implemented the mesh distribution in the ANSYS software tool with different element sizes in order to perform the numerical modeling of 8/6 SRM. The size of the elements is very important for the calculation of the global electromagnetic quantities. Smaller items give more accurate results and longer computation time, and vice versa. We observe that the technical meshes used in this modelization are the following: the elements large within the limits of the machine and getting smaller and smaller at the air gap where the line of sliding, conversely, more and more in the center (The tree) of the SRM. The technical mesh used in Figure 2 allows exploiting the computational time and for accurate results, the smaller items are used just in regions where variations in the values of the results are significant. The most important elements are adopted in regions where these variations are not significant. This economic and technical choice of elements distribution model will be used throughout the modeling of our study either linear or non-linear case. This choice ensures an important very good quality of results. The static characteristics of the torque and the inductance in SRM will be determined by static modeling in FEM which has been effective [13-14], [16], where it can be taken into account magnetic saturation effects and air gap variation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002611_j.promfg.2020.04.165-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002611_j.promfg.2020.04.165-Figure2-1.png", + "caption": "Fig. 2 a) Printed preform, b) Forged part", + "texts": [ + "27mm thick PA12/carbon fibers produce preforms with a theoretical final thickness of 18.9mm. The printing temperature is set at 200\u00b0C and the feed ratio at 4mm/s. Fibres are oriented around the part. Table 1 shows the characteristics provided by 9T Labs for the consolidated composite yarn. Table. 1: Material datasheet CF/PA 12 9T LABS (credit 9T Labs). Carbon fibre STS 40 F13 (24k) Toho Tenax Matrix system PA12 EMS chemie Fibre volume content 50% Density of composite 1.405g/cm3 Forging step The forging step consists of: The placement of the preform with an initial height \u210e\ud835\udc56\ud835\udc56 (Figure 2.a)) in a steel mold. Heats the tools and the preform with a convection oven up to a temperature of 250\u00b0C until the temperature in the composite, controlled by thermocouples, reach the set. Compression of the part with an Instron traction/compression machine. Pressure is applied in Z direction (direction of \u210e\ud835\udc56\ud835\udc56 and \u210e\ud835\udc53\ud835\udc53 on Fig. 2) until a given height \u210e\ud835\udc53\ud835\udc53. Material mainly moves in the Z direction but also in the other directions so as to fill the functional gaps of tool. The compression speed is set to 10mm/min and the compression force is measured during the process. Demolding of the part. Deburring: mold is not fully closed, if a high compression ratio is applied this can cause a burr. Five parts with different compression ratios were produced. Figure 2.a) shows a preform obtained by 3D printing. This preform is rigid and easy to handle. Figure 2.b) shows a part obtained after forging, the final height depends on the compression ratio applied. Parts characteristics Parts are weighted with a 0.1 mg accuracy after printing and on the final part (after deburring). The material loss cause by deburring \ud835\udf0f\ud835\udf0f\ud835\udc59\ud835\udc59 is then determined by the equation (1) where \ud835\udc5a\ud835\udc5a\ud835\udc56\ud835\udc56 and \ud835\udc5a\ud835\udc5a\ud835\udc53\ud835\udc53 are respectively the mass of the printed and final parts. Victor Haguenauer et al. / Procedia Manufacturing 47 (2020) 169\u2013173 171 Author name / Procedia Manufacturing 00 (2019) 000\u2013000 [7] studied the impact of porosities on the limit values at rupture of composite plates", + "27mm thick PA12/carbon fibers produce preforms with a theoretical final thickness of 18.9mm. The printing temperature is set at 200\u00b0C and the feed ratio at 4mm/s. Fibres are oriented around the part. Table 1 shows the characteristics provided by 9T Labs for the consolidated composite yarn. Table. 1: Material datasheet CF/PA 12 9T LABS (credit 9T Labs). Carbon fibre STS 40 F13 (24k) Toho Tenax Matrix system PA12 EMS chemie Fibre volume content 50% Density of composite 1.405g/cm3 Forging step The forging step consists of: The placement of the preform with an initial height \u210e\ud835\udc56\ud835\udc56 (Figure 2.a)) in a steel mold. Heats the tools and the preform with a convection oven up to a temperature of 250\u00b0C until the temperature in the composite, controlled by thermocouples, reach the set. Compression of the part with an Instron traction/compression machine. Pressure is applied in Z direction (direction of \u210e\ud835\udc56\ud835\udc56 and \u210e\ud835\udc53\ud835\udc53 on Fig. 2) until a given height \u210e\ud835\udc53\ud835\udc53. Material mainly moves in the Z direction but also in the other directions so as to fill the functional gaps of tool. The compression speed is set to 10mm/min and the compression force is measured during the process. Demolding of the part. Deburring: mold is not fully closed, if a high compression ratio is applied this can cause a burr. Fig. 1 CarbonKit 9T Labs CF composite long fibre printing system implanted on a Ultimaker 3D printer (credit 9tlabs) Five parts with different compression ratios were produced. Figure 2.a) shows a preform obtained by 3D printing. This preform is rigid and easy to handle. Figure 2.b) shows a part obtained after forging, the final height depends on the compression ratio applied. Fig. 2 a) Printed preform, b) Forged part 2.2. Sampling and analysis Parts characteristics Parts are weighted with a 0.1 mg accuracy after printing and on the final part (after deburring). The material loss cause by deburring \ud835\udf0f\ud835\udf0f\ud835\udc59\ud835\udc59 is then determined by the equation (1) where \ud835\udc5a\ud835\udc5a\ud835\udc56\ud835\udc56 and \ud835\udc5a\ud835\udc5a\ud835\udc53\ud835\udc53 are respectively the mass of the printed and final parts. Author name / Procedia Manufacturing 00 (2019) 000\u2013000 3 \ud835\udf0f\ud835\udf0f\ud835\udc59\ud835\udc59 = 1 \u2212 \ud835\udc5a\ud835\udc5a\ud835\udc53\ud835\udc53 \ud835\udc5a\ud835\udc5a\ud835\udc56\ud835\udc56 (1) Dimensions of the manufactured parts are also measured. The dimensions before and after shaping make it possible to define the compression ratio from the initial height \u210e\ud835\udc56\ud835\udc56 and final height \u210e\ud835\udc53\ud835\udc53 of the part: \ud835\udf0f\ud835\udf0f\ud835\udc50\ud835\udc50 = 1 \u2212 \u210e\ud835\udc53\ud835\udc53 \u210e\ud835\udc56\ud835\udc56 (2) Micro and macrographic analysis After shaping, the parts produced are sectioned in order to produce samples which can be used for carrying out micrographs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000609_978-3-030-12684-1_17-Figure17.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000609_978-3-030-12684-1_17-Figure17.2-1.png", + "caption": "Fig. 17.2 Three unique build orientations", + "texts": [ + " Test specimens used for the series of experiments herein are built with ABSplus-P340 polymer using a uPrint SE Plus [30], or with 316L stainless steel using a Concept Laser M2, which is a direct metal laser sintering machine [31]. The parts are designed to be rectangular cantilevers with high aspect ratio cross-sections to obtain relatively large displacements normal to the largest face and provide a wide, flat surface convenient for DIC speckling. Three solid ABS parts are printed with the dimensions shown in Fig. 17.1. Each part is built in a different orientation, as shown in Fig. 17.2. That is, all three parts in this set have the same outer geometry, but have their build layers oriented in different Cartesian directions. A second set of three ABS parts is printed with internal lattices. Each lattice is a 2D cell pattern printed along the build orientation, as shown in Fig. 17.3. Thus the second set of parts corresponds to the same build orientations as in Fig. 17.2, with lattices oriented vertically with respect to the base plate. The lattices are only built in the cantilever section of each part\u2014the base is left solid. The 2D cell pattern consists of 1.5 mm square holes separated by 1 mm solid walls, and surrounded by a 1 mm thick external wall. Figure 17.4 shows the outer dimensions of the lattice ABS set, which are increased from that of the solid ABS parts to allow for the internal lattice structures. A third set of parts is built in solid steel. These parts are built with the same geometry as the solid ABS parts (Fig. 17.1), except they are 3 mm thick instead of 4 mm, and have a 5 mm fillet instead of 4 mm. The three steel parts are also built in the same orientations as the three solid ABS parts (Fig. 17.2). The parts were excited by the translational oscillations of a shaker table using the Shaker Control software from Bruela & Kjaer Sound and Vibration Measurement. The shaker table was oriented so that the large face of the part oscillated normal to the imaging plane. Parts are secured to the shaker table with 4 mm bolts, as seen in Fig. 17.5. The bolts are torqued to 1.75 Nm. Parts were tested with a random excitation over select frequency ranges. Measurements for out of plane modes were collected in the following ranges: mode 1 at 20\u2013100 Hz, etc" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000827_1350650119866040-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000827_1350650119866040-Figure2-1.png", + "caption": "Figure 2. Axle bearing assembly with two cylindrical roller bearings mounted on the journal of wheel shaft of railroad cars: (a) axle bearings assembly system structure, (b) multi-interfacial contact mechanics model, (c) location number of rollers of bearings and radial load distribution and (d) typical damage to the inner race of the inboard bearing.", + "texts": [ + " Between November 2005 and August 2009, there were a total of 40 early flaking failures on bearings NJ3226 and NJP3226 of the axle bearings assembly in some areas of China\u2019s railway. The running time of the cars assembled with these bearings were less than two years (2.5 years required), and the ratio of early fatigue in the same bearings was 0.2%; NJ3226 was the main early failure bearing, accounting for 96% of the total number of failed bearings and fatigue proportion of its inner race was 87%, especially occurred near the rib.1 The typical damage to the inner race of the inboard bearing is shown in Figure 2(d).1 This paper attempts to explain this actual engineering problem using a systematic methodology. Even though many bearing failures of railroad car could be found, often caused by surface defects, corrosion, broken axle and structural irrational assembly in the actual vehicle operation,2,3 over last decades, the similar investigation of axle bearings assembly with cylindrical roller bearings (CRBs) was relatively limited. However, several basic theories had been established on numerical methods,4 contact,5,6 lubrication,7\u201311 service life12\u201314 of CRBs by many early scientists on roller bearings", + " In the installation process of CRBs assembly, the interference fit is applied between the bearings and the axle journal, and the clearance fit is applied between the bearings and the housing. Two sets of bearings NJ3226 and NJP3226 have different structure design and installation requirement. Components of the whole axle bearings assembly includes: (1) front axle box cover, (2) bolts, (3) resilient pad, (4) end cup, (5) flat ring, (6) outer ring of NJP3226, (7) cage, (8) roller, (9) inner ring of NJP3226, (10) outer ring of NJ3226, (11) axle box, (12) inner ring of NJ3226, (13) metal labyrinth seal, and (14) journal of shaft (type number of RD3), as shown in Figure 2(a), and its design parameters are as shown in Table 1. NJ3226 near the wheel and NJP3226 away from the wheel are, respectively, referred as \u2020, I\u2020 in Figure 2(b). Under given load and material properties, the real contact performance depends largely on clearance, interference fit of parts, and fit differences between them. These design parameters are controlled in a certain range from assembly manual and industry standards in Table 2. Location number of rollers and radial load distribution in this case are shown in Figure 2(c). Full numerical finite element model of axle bearings assembly considering multi-interface contact is established as shown in Figure 3, considering the internal radial and axial clearances of bearings, the interference fit between inner ring and journal as well as the clearance between outer ring and axle box housing. The model contains roughly three million of mesh grids with the cantilever beam structure. Multi-interface contact mechanics model is employed to study the contact load distribution and deformation of the entire system under actual working conditions, in which the operating radial load applied on the axle housing is 89 kN calculated by UIC510-5 (standard of International Union of Railways) and related parameters as shown in Table 1. In addition, the surface load that applied on the surface of the load area of the bearing housing19 equals the operating load as shown in Figure 2(c). The results had been discussed in Guo et al.19 and Chang et al.19,20 and the inner contact pressure can be calculated by FEM. However, in order to analyze the local contact and deformation between roller and raceway more accurately and evaluate the subsequent lubrication performance and fatigue life of bearings, the classical non-Hertzian contact numerical model is used. In this paper, the misalignment angles and load distribution calculated from FEM as initial values are used by models below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003377_ccdc49329.2020.9164095-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003377_ccdc49329.2020.9164095-Figure3-1.png", + "caption": "Figure 3: Flight Configuration of SAW(\u03b4 = \u221230\u25e6)", + "texts": [], + "surrounding_texts": [ + "d\u03072k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)k(ld sin \u03b4k)\u03b4\u0307k \u2212(l cos \u03b4k)\u03b4\u0307k\n\u23a4 \u23a6 ,\nand\nd\u03082k\u22121(ld) =\n\u23a1 \u23a3\n0 0 0\n\u23a4 \u23a6 ,\nd\u03082k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)kld(cos \u03b4k \u03b4\u0307 2 k + sin \u03b4k \u03b4\u0308k)\nld(sin \u03b4k \u03b4\u0307 2 k \u2212 cos \u03b4k \u03b4\u0308k)\n\u23a4 \u23a6 .\nBecause the first derivative of the \u03b4k represents the folding angular rate, and in the assumptions presented above, the folding angular rate during morphing process keeps constant, so the second derivative of \u03b4k is zero, thus the second derivative of d2k can be expressed as follows:\nd\u03082k(ld) =\n\u23a1 \u23a3\n0\n(\u22121)kld(cos \u03b4k)\u03b4\u0307 2 k\n(ld sin \u03b4k)\u03b4\u0307 2 k\n\u23a4 \u23a6 . (6)\nAfter computing d\u0307 and d\u0308 in equation (2) and (3) above, Meanwhile, the domain item dcm of Fext in (2) can also be decomposed in two parts:\nd\u0308cm = d\u0308cm1 + d\u0308cm2, (7)\nand\ndcm1 = 1\n2m mf lf (R1 +R3),\ndcm2 = 1\n2m mtipltip(R2 +R4),\n(8)\nwhere Ri(i = 1, 2, 3, 4) are represented as follows:\nR1 = 0, R2 = \u03b4\u030721 [0,\u2212 cos \u03b41, sin \u03b41] T , R3 = 0, R4 = \u03b4\u030722 [0, cos \u03b42, sin \u03b42] T ,\nSimilarly, the additional moment caused by morphing of wing tips can be computed. The domain item ( \u222b [d\u0304]d\u0308dm) in Mext can be given as follows: \u222b\n[d\u0304]d\u0308dm = 1\n2 [mf lf (S1 + S3)\n+ mtipltip(S2 + S4)][1, 0, 0] T , (9)\nwhere Si(i = 1, 2, 3, 4) represent the additional terms caused by folding of wing tip:\nS1 = 0, S2 = (l0 \u2212 ltip)\u03b4\u0307 2 1 sin \u03b41,\nS3 = 0, S4 = (ltip \u2212 l0)\u03b4\u0307 2 2 sin \u03b42,\nWhere l0 is the half width of the front view fuselage. Then, substituting the domain item d\u0308cm in Fext and\nthe domain item \u222b [d\u0304]d\u0308dm in Mext into the equation above, the Fext and Mext can be given as follows:\nFext = \u22121\n2 mtipltip(R\u03b41\u0394\u03b41 +R\u03b42\u0394\u03b42), (10)\nwhere R\u03b41 = \u03b4\u030721 [0, sin \u03b41, cos \u03b41] T ,\nR\u03b42 = \u03b4\u030722 [0,\u2212 sin \u03b42, cos \u03b42] T ,\nand\nMext = \u22121\n2 mtipltip(S\u03b41\u0394\u03b41 + S\u03b42\u0394\u03b42)[1, 0, 0]\nT , (11)\nwhere S\u03b41 = (l0 \u2212 ltip)\u03b4\u0307 2 1 cos \u03b41,\nS\u03b42 = (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42.\nThen, substituting equation (10) and (11) into equation (2) and (3), the nonlinear model of the folding wing-tip UAV can be expressed as follows:\nu\u0307 = rv + qw \u2212 g sin \u03b8 + Fx\nm , (12)\nv\u0307 = \u2212ur + wp+ gcos \u03b8sin\u03c6+ Fy\nm\n\u2212 mtipltip 2m (\u03b4\u030721 sin \u03b41\u0394\u03b41 \u2212 \u03b4\u030722 sin \u03b42\u0394\u03b42),\n(13)\nw\u0307 = uq \u2212 vp+ gcos \u03b8cos\u03c6+ Fz\nm\n\u2212 mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42),\n(14)\np\u0307 = (c0r + c1p)q + c2L\u0304+ c3N \u2212 c4 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42),\n(15)\nq\u0307 = c5pr \u2212 c6(p 2 \u2212 r2) + c7M, (16)\nr\u0307 = (c8p\u2212 c2r)q + c4L\u0304+ c9N, (17)\nwhere \u03b4\u03071,2 is the folding rate of the wing tip, \u0394\u03b41,2 is the folding angle change in a period of folding process, L\u0304,M ,N are components of total moment along three axes within body frame and can be given as follows:\nL\u0304 =p\u0307Ix \u2212 r\u0307Ixz + qr(Iz \u2212 Iy)\u2212 pqIxz \u2212 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42)\nM = Iy q\u0307 + pr(Ix \u2212 Iz) + (p2 \u2212 r2)Ixz\nN = r\u0307Iz \u2212 p\u0307Ixz + pq(Iy \u2212 Ix) + qrIxz\nwhere ci(i = 0, 1, 2, ..., 9) are constant coefficients expressed as follows:\nc0 = ( (Iy \u2212 Iz)Iz \u2212 I2xz\nIxIz \u2212 I2xz ), c1 = (Ix \u2212 Iy + Iz)Iz \u2212 Ixz IxIz \u2212 I2xz\nc2 = Iz\nIxIz \u2212 I2xz , c3 = Iz IxIz \u2212 I2xz , c4 = 1 IxIz \u2212 I2xz\nc5 = Iz \u2212 Ix\nIy , c6 = Ixz Iy , c7 = 1 Iy ,\nc8 = Ix(Ix \u2212 Iy) + I2xz\nIxIz \u2212 I2xz , c9 = Ix IxIz \u2212 I2xz\n1816 2020 Chinese Control And Decision Conference (CCDC 2020)\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply.", + "And Fx,Fy ,Fz are components of aerodynamic force and propulsive force along three axes within body frame, given that the proplusive force is along body x-axis and the engine offset angle \u03b1T = \u03b2T = 0, namely, T = Tx. Based on the conversion relationship between the body frame and airflow frame Xbody = ST \u03b1\u03b2Xwind, there exists:\n\u23a1 \u23a3 Fx\nFy Fz \u23a4 \u23a6 body = \u23a1 \u23a3 T 0 0 \u23a4 \u23a6 body + ST \u03b1\u03b2 \u23a1 \u23a3 \u2212D Y \u2212L \u23a4 \u23a6 wind (18)\nHence, Fx,Fy ,Fz are given as follows:\nFx=T + L sin\u03b1\u2212 Y cos\u03b1 sin\u03b2 \u2212D cos\u03b1 cos\u03b2, (19)\nFy = Y cos\u03b2 \u2212D sin\u03b2, (20)\nFz=\u2212L cos\u03b1\u2212 Y sin\u03b1 sin\u03b2 \u2212D sin\u03b1 cos\u03b2, (21)\nwhere L,Y ,D are thrust, side and drag force. \u03b1,\u03b2 are angleof-attack and the sideslip angle, \u03b8 and \u03c6 are pitch angle and roll angle.\nWith decoupling method proposed in [21], the decoupled longitudinal nonlinear model is given as follows:\nmV\u0307 = T cos\u03b1\u2212D+mg(\u2212cos\u03b1 sin \u03b8+sin\u03b1 cos \u03b8), (22)\nmV \u03bc\u0307 = T sin\u03b1+ L\u2212mg(sin\u03b1 sin \u03b8 + cos\u03b1 cos \u03b8)\n\u2212mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (23)\nmV \u03b1\u0307 = \u2212T sin\u03b1\u2212L+mV q+mg(sin\u03b1 sin \u03b8+cos\u03b1 cos \u03b8)\n\u2212mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (24)\n\u03b8\u0307 = q, (25)\nIy q\u0307 = M, (26)\nSimilarly, the decoupled lateral nonlinear model is given as follows:\nmV \u03b2\u0307 = Y \u2212mV (\u2212p sin\u03b1) + r cos\u03b1\n\u2212 mtipltip 2m (\u03b4\u030721 cos \u03b41\u0394\u03b41 + \u03b4\u030722 cos \u03b42\u0394\u03b42), (27)\n\u03c6\u0307 = p+ (r cos\u03c6+ q sin\u03c6) tan \u03b8, (28)\n\u03c8\u0307 = 1\ncos \u03b8 (r cos\u03c6+ q sin\u03c6), (29)\np\u0307 = (c0r + c1p)q + c2L\u0304+ c3N \u2212 c4 mtip\n2m ltip((l0\n\u2212 ltip)\u03b4\u0307 2 1 cos \u03b41\u0394\u03b41 + (ltip \u2212 l0)\u03b4\u0307 2 2 cos \u03b42\u0394\u03b42),\n(30)\nr\u0307 = (c8p\u2212 c2r)q + c4L\u0304+ c9N, (31)\nIt can be obtained from the equation (12)-(17) that during the morphing process of folding wing-tip UAV, some aerodynamic items such as the x-axis direction component of airspeed V , pitch and yaw angular rate, seem to have no distinction with that of conventional aircrafts. Consequently, It can be derived from above that the whole morphing process, including taking off, taking up and cruise, only has impacts on the aerodynamic performance of the folding wing-tip UAV in y-axis or z-axis, and in x-axis, the UAV appears the same as conventional aircrafts.\n3 Numerical Simulation\nTo validate the utility of the modeling method and nonlinear models of the folding wing-tip UAV, a numerical study is conducted to SAW, which is a typical kind of folding wing-tip UAV. The 3D model of the SAW is established in DATCOM and aerodynamic performances under different folding angles are analyzed based on numerical simulation. The basic airframe parameters of SAW are listed as follows:\nBased on above airframe parameters, with the application of DATCOM, the 3D model of the SAW can be established and the flight status(\u03b4 = \u221230\u25e6 and \u03b4 = 60\u25e6) can be shown as follows:\n2020 Chinese Control And Decision Conference (CCDC 2020) 1817\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply.", + "With DATCOM symbols and the corresponding shape parameter given above, the related parameters in DATCOM such as SSPNE and SSPN corresponding to different folding angles can be obtained and listed below, based on which numeric simulation can be conducted to obtain aerodynamic performance of SAW under different folding angles:\nThus, substituting all the parameters and symbol values into DATCOM, the aerodynamic coefficients with folding angles ranging from \u221230\u25e6 to 60\u25e6 as well as control surface items can be obtained. Based on this, the aerodynamic performances under different folding angles with angle of attack ranging from \u22124\u25e6 to 10\u25e6 can be shown below:\nIt can be illustrated in Figure 5 that during morphing process, the lift coefficients appear linear correlation with angle of attack, meanwhile, it appears that with folding angle keeping at \u221260\u25e6, the lift coefficients have a greater slope, which shows folding upwards can make it easier to enhance altitude when the SAW takes off. Similarly, in Figure 6, with folding angle keeps at 30\u25e6, the drag coefficients are much smaller, which make it more suitable in cruise phase.\nTo clearly clarify the ability of SAW to conduct multimissions, an insightful description of the polar curves of the SAW under different folding angles are given:\nIn Figure 8, when \u03b4 = 0\u25e6, the lift-to-drag ratio reaches the maximum value, which means that keeping the wing-tip level is more suitable to conduct long-range cruise surveillance missions. Folding wing-tip leads to the decrease of lift-to-drag ratio, which makes the SAW have the high-speed airfoil with small-aspect-ratio and large-sweep-angle. With the wing-tip folding upwards, the SAW can dive fast and conduct high-speed sprint, which has a greater maneuverability.\n1818 2020 Chinese Control And Decision Conference (CCDC 2020)\nAuthorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 07:29:26 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_80_0001549_s11223-019-00110-8-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001549_s11223-019-00110-8-Figure3-1.png", + "caption": "Fig. 3. Force analysis diagram of the movable tooth.", + "texts": [ + " The speed of B in the tangent direction on the movable tooth is the sliding speed of the movable tooth relative to the ring gear: V B t AO B OV r 1 1 1cos (cos ) sin ( ) cos( B G H G H H G Hc l k Z Z ckZ1 2 2 2 2 2 2 2 2 2 A HZ)sin( ), (4) where A is the included angle of V AO1 and VA , RA is the polar diameter at the point A of the actual contour of the wave generator, O1 is the angle between V O1 and the x-axis, H is the angle between the swing-rod and the x-axis, A is the argument of point A on the actual contour of the wave generator, and ZH is the teeth number of the wave generator, ZH 2. By analyzing the force balance condition of a movable tooth, it can be seen that the deformation of a movable tooth along the normal direction of the contact area in the working area of a wave generator is distributed according to the sine law (Fig. 3). Forces FHi and exerted on the movable tooth by the wave generator and ring gear are F M Z Hi i 135 2. sin( ), (5) F M Z Ki i i i i i 135 2 2 2 1 2 . sin( )cos( ) cos( ) . (6) The contact stress of the movable tooth and ring gear can be inferred from the elastic Hertz formula as K E i i i d i i Z M Z L 135 2 2 2 1 2 . sin( )cos( ) cos( ) , (7) where d is the equivalent radius of curvature, L is the working width of the movable tooth, i is the angle between OOi and the tangent line of ki the ring gear at point A, 1i is the angle between OOi and FHi , Z is constant, 2i is the angle between OOi and FKi , and i is the angle between OOi and the x-axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000436_wits.2019.8723658-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000436_wits.2019.8723658-Figure2-1.png", + "caption": "Fig. 2. Air motor type palette ATLAS COPCO LZB14", + "texts": [], + "surrounding_texts": [ + "In this section the modeling of the pneumatic motor is important to observe the pneumatic energy conversion and then to simulate and evaluate the strategies proposed , The air motor presented in these study : air motor ATLAS COPCO LZB14 AR00511, This motor presents the lot of advantages such as : safety, simplicity, lightness and high power density that make it suitable for portable applications. [2] [12] [11]" + ] + }, + { + "image_filename": "designv11_80_0002154_chilecon47746.2019.8986859-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002154_chilecon47746.2019.8986859-Figure1-1.png", + "caption": "Fig. 1 Sketch of a possible scenario where a nano robot (left side cylinders) exerts an electrical force on the virus by locate them away from these artificial cells. )", + "texts": [ + " Normally, when the electric potential is known one would take the well-known Poisson\u2019s equation ~\u22072\u03a6(~r) = 4\u03c0\u03b5\u03c1 where \u03a6 denotes the electric potential in the point ~r. So often one needs to translate the problem of the solution of the Poisson\u2019s equation to a well-defined coordinates system. So now we will write the following assumption based in experimental evidence: (i) Bacterium posses a well defines geometrical form which might be well approximate to a spatial geometry. Therefore we consider the scenario by which a bacterium has a curved geometry as the one sketched in Fig.1. (ii) Because the ATP processes most of the charged ions would lie inside the virus body, so that aggregation of the positive ions inside of the virus would constitute a charge density so exists inside electric charge inside of them. (iii) Due to the ATP processes that keeps continuously the homeostasis of the internal electrical processes in bacterium is assumed that the charges or ions of Potassium mostly in a static state fact that does not allow the creation of internal currents or variations of the charge density in time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002077_iros40897.2019.8968281-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002077_iros40897.2019.8968281-Figure2-1.png", + "caption": "Figure 2. The first structure of the helical actuator.", + "texts": [ + " A typical example of the working performance of the helical actuator is shown in Fig. 1(b), where a helical actuator with 11.6mm outer diameter tied with 200 g load coils under 0.3 MPa air pressure. The shrinkage length is 60% of the initial length of the actuator, meaning 60% contraction ratio is realized. 978-1-7281-4004-9/19/$31.00 \u00a92019 IEEE 8300 Based on different application circumstances, we propose 2 kinds of structures and fabrication methods for the helical actuator. The first kind of structure of the helical actuator is shown in Fig. 2. The one-way extensible cloth refers to the cloth made by braiding elastic string and non-stretchable fiber perpendicularly so that one direction is extensible while the vertical direction is not. The fiber angle means the acute angle between fiber\u2019s weaving direction (inextensible direction) and the axial direction of the tube, this angle influences the shape of the helical actuator after pressurized. As it is shown in Fig. 2, the fabrication process which requires rubber tube, one-way extensible cloth, and inextensible wire can be conducted as follows: 1. Place the inextensible wire inside the tube on one side of the inner surface and place the tube on a piece of one-way extensible cloth. 2. Sew the cloth together into a sleeve along the tube. 3. Fix inlet and block respectively at 2 ends of the tube and fix the wire, tube, cloth and inlet (block) together, ensure the wire is attached on the same side of the inner surface of the tube" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002029_iccas47443.2019.8971518-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002029_iccas47443.2019.8971518-Figure5-1.png", + "caption": "Fig. 5 Folding and spreading motion of forelimbs.", + "texts": [ + " Using the rotating motor attached to the upper part, the rack moves forward and backward through the movement of the crank shaft and the worm gear, and the force is distributed and transmitted to the lower part having the same structure by using connecting the belt. In addition, a linear actuator and a link structure are designed to mimic the mole scapula and anatomy on both sides. Synthetically, the forelimbs move forward through one rotary motor and two linear actuators and make the motion that gathering paws to the center, pushing arms to both sides and returning to the initial position. The removal of excavated soil proceeds through the repetition of this motion. Fig. 5 Folding and spreading motion of forelimbs. Figure 6 shows the overall design combined expandable drill bit and forelimbs. Drill bit module keeps advanced state while drilling, and when debris removal is started, it moves backward into the body, so that interference between the drill bit and forelimbs can be prevented. Fig. 6 Overall structure of drill bit and forelimbs. The sequence of the whole drilling mechanism of the proposed robot is as follows. 1) Advance drill bit module, 2) Rotate while expanding blades, 3) Excavation, 4) Contraction blades and reverse drill bit module, 5) Move forward and folding of forelimbs, 6) Remove debris spreading and move backward forelimbs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002025_iccas47443.2019.8971672-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002025_iccas47443.2019.8971672-Figure6-1.png", + "caption": "Fig. 6 An appearance of the proposed training walker. It has an upper extremities load measuring device and a dual-arm type gait handling device.", + "texts": [ + " Figure 5 shows the maximum and minimum assisting force/moment applied by the PT to one subject in a gate experiment. From the results, it can be seen that the PT applied a maximum of -13N in the X direction and a maximum of 20N in the Z direction, which means that he lifted the subject\u2019s COM while pulling it backward. As for a moment, the PT applied a maximum of -2Nm around the X axis and a maximum of 5Nm around the Z axis, which means that he rotated the subject\u2019s pelvic while obliquing it. These results can be usefully referenced in designing the pelvic handling device. Figure 6 shows the appearance of the proposed training walker and its mechanical configuration. The walker is constructed based on a commercially available standup rolling walker (Paramount Co.Ltd., KA-391). Two distinctive modules are implemented in the walker; one is a device for measuring an upper extremities load, and the other is a dual arm-type gait handling device. The load measuring device is mounted under the armrest, which has four load sensors that can measure the upper extremities load. The measured load information is used to evaluate the patient\u2019s gait ability and to control the gate handling device to effectively handle a patient\u2019s gait", + " As for a moment, the PT applied a maximum of -2Nm around the X axis and a maximum of 5Nm around the Z axis, which means that he rotated the subject\u2019s pelvic while obliquing it. These results can be usefully referenced in designing the pelvic handling device. Fig. 5 Experimental results of the assist (a) force and (b) moment by a PT. The PT applied a maximum of -13N (X), 20N (Z), 2Nm (around X axis), and 5Nm (around Z axis). Upper-limb load measuring equipment Pelvic handling arm \u00b145deg \u00b150mm \u00b145deg 6 8 0 m m 520mm Load sensor J0 J1 J2 J0 J1 J2 Fig. 6 An appearance of the proposed training walker. It has an upper extremities load measuring device and a dual-arm type gait handling device. 4. TRAININGWALKERWITH GAIT HANDLING ARM 4.1 Training walker with a gait handling arm Figure 6 shows the appearance of the proposed training walker and its mechanical configuration. The walker is constructed based on a commercially available standup rolling walker (Paramount Co.Ltd., KA-391). Two distinctive modules are implemented in the walker; one is a device for measuring an upper extremities load, and the other is a dual arm-type gait handling device. The load measuring device is mounted under the armrest, which has four load sensors that can measure the upper extremities load. The measured load information is used to evaluate the patient\u2019s gait ability and to control the gate handling device to effectively handle a patient\u2019s gait" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002959_052080-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002959_052080-Figure5-1.png", + "caption": "Figure 5. Prototype Component Connection of Full Circuit", + "texts": [ + " So, the single pole single through switch was representing the PP float switch. The Arduino UNO operated depend on Arduino coding. Arduino coding use Arduino IDE to write the code and upload the coding file to Arduino UNO at Proteus 8 Professional to simulate system operation JICETS 2019 Journal of Physics: Conference Series 1529 (2020) 052080 IOP Publishing doi:10.1088/1742-6596/1529/5/052080 After that the real component was develop as designed with PV system to supply the power to system. The component connection complete irrigation system shown in Figure 5. This section explains the result of the project. It covered result of the simulation and hardware of automated irrigation system. Simulation circuit consists of the designed circuit of the system configuration. The status of water level detected by water level sensor and the controller switched the water pump to ON and OFF based on condition state at coding Arduino IDE. When low water level detected, PP float switch ON and sent signal to controller. The controller delivered the signal to run the water pump" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001699_1350650119893896-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001699_1350650119893896-Figure10-1.png", + "caption": "Figure 10. Nonlinear stability boundary and orbits corresponding to different shaft rotating speed: (a) ! \u00bc 2:305, (b) ! \u00bc 2:372, (c) ! \u00bc 2:437, and (d) ! \u00bc 2:5.", + "texts": [ + " Repeat the process of selecting middle point in the updated line OSO1 or O1Occ until the distance between On 1 and On is small enough in nth iteration, then On could be considered as the critical point Ocritical on line OSOcc. 4. Repeat the above process until the 24 critical points are determined, then connect the 24 critical points with a line and the stability boundary can be obtained. With the identical external load but different rotating speeds, the nonlinear stability boundaries of finitelength hole-entry hybrid journal bearings are shown in Figure 10. The orbits corresponding to different initial points are also given. The working conditions are the same as listed in Table 1. When ! \u00bc 2:305, the orbits can be stable in the entire bearing clearance as shown in Figure 10(a). In Figure 10(b) to (d), with an increase of the shaft rotating speed, the nonlinear stability boundary is becoming smaller. The reason is that with higher rotating speed, the inertial forces of the oil become stronger and the oil film is more unstable. The shape of the boundary is quasi circle with the equilibrium point Os as the center. If the rotating speed goes up to the threshold speed, the nonlinear stability boundary will shrink to the equilibrium point Os. The differences between linear and nonlinear stabilities are summarized as the color band in Figure 11", + " However, under nonlinear conditions, even the shaft rotating speed is smaller than threshold speed, there exists an obvious transitional region. When the rotating speed is lower than the transitional speed 05 !5 !trans\u00f0 \u00de, the nonlinear stability boundary almost coincides with the clearance circle, which means the shaft center can always return to the stable equilibrium point. When the rotating speed is increased to the transitional region !trans 5 !5 !th\u00f0 \u00de, the stability of the bearing system relies on the initial point of the shaft center. The stable region becomes smaller with a higher rotating speed as shown in Figure 10. When rotating speed exceeds the threshold speed !4 !th\u00f0 \u00de, a slight fluctuation can introduce instability to the bearing system. As compared to the nonlinear assumption, the linear one enlarges the stability region. In practical engineering, the shaft is likely to suffer many disturbances. If the threshold of the rotating speed is calculated from the linear assumption, the bearing system with rotating speed in the transitional region could unexpectedly encounter the instability. To ensure the stability of operation, the system should be analyzed under nonlinear conditions, and the rotating speed should be lower the transitional speed " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003106_mrc.2019.153-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003106_mrc.2019.153-Figure1-1.png", + "caption": "Figure 1. (a) The CAD design and geometry of the hybrid structure consists of a space-filling truss alumina nanolattice comprised of hollow tubes sandwiched between two alumina plates. (b) The unit cell is comprised of hollow tubes and exhibits a cubic symmetry. (c) The hollow tube.", + "texts": [ + " We found that in addition to the usual scaling laws present in the nanolattice structures, there were also unusual scaling relations between stiffness/strength and density for these lattice-plate hybrid structures. These unusual scaling relations likely originate from the complex stress states at the joints between the plates and the nanolattices. Hence, the present work and its findings are able to provide guidelines for the design of advanced architectured materials. Structure design, material, and finite element model Figure 1 shows the architectures of the hybrid structure, which consists of a space-filling truss alumina nanolattice as a core and two alumina plates as face-sheets, and the details of the hybrid structure geometry are also given in the figure. To reduce the weight, the nanolattice is comprised of hollow alumina tubes. Many previous studies[22\u201324] have revealed that highly periodic structures are able to give high mechanical stability and stiffness. Therefore, optimal results can be expected by choosing different unit cells for the lattices" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002139_s11223-020-00139-0-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002139_s11223-020-00139-0-Figure5-1.png", + "caption": "Fig. 5. Histograms of pore (a) size distribution and fractal dimension (b).", + "texts": [ + " The size of pores determines the value of SIF, and the coefficient Y takes into account the geometrical dimensions of the specimen and loading conditions. The fractal dimension FD characterizes the value of stress concentration as a pore shape function. The results of an analysis of the pore space of investigated sintered specimens allowed us to determine the regularities of the distribution of the morphology characteristics of pores and their average values taking into account the distribution law (Fig. 5). It can be seen that the pore size distribution follows well the normal distribution law. This indicates the random nature of the observed quantity. At the same time, the values of the fractal dimension of pores are in poorer agreement with normal distribution, indicating that they are affected by extraneous factors. In the case of synthesis of alloys based on PT5 titanium powder, the pore shape is probably influenced by the peculiarities of the morphology of particles of powder obtained by the mechanical milling of a titanium sponge [25]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001442_042023-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001442_042023-Figure2-1.png", + "caption": "Fig 2. The schematic of suspension part", + "texts": [], + "surrounding_texts": [ + "At present, the method of \u201cinsulated rope twisting\u201d is widely utilized in the removal of foreign objects in power lines. This method uses an insulating rope with a middle bolt spring-like wire hook to throw the wire, and the grounding worker pulls the insulating rope to the foreign object, and the twisting of the spring-shaped wire hook causes the foreign objects to be twisted with the rope. Then the foreign objects can be removed by pulling on the rope. The method of \u201cinsulated rope twisting\u201d has the following drawbacks. Firstly, in the case of solid foreign objects, this method often causes foreign objects and insulated rope hooks to entangle the wires, which make it difficult to remove foreign objects. Secondly, the wire hooks directly contact the lines. If the force is too vigorous, then the insulation performance of the insulated rope is decreased. Thirdly, for this method, it is difficult to control the force. Therefore, the use of this method may cause line oscillation and phase-to-phase short circuit, and also jeopardizes the safety of operator [3]. In this paper, a new foreign objects removal device used in high-voltage transmission line is proposed, which overcomes the shortcomings of traditional devices. The proposed foreign objects removal device is introduced in the next sections." + ] + }, + { + "image_filename": "designv11_80_0001441_012149-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001441_012149-Figure10-1.png", + "caption": "Figure 10. Integrated variable speed moto-compressor.", + "texts": [ + " The second choice is to optimize the speed and to cope with the volume flow of each compressor minimizing the weight and the footprint and maximizing the efficiency of each compressor, introducing the use of centrifugal compressors for high pressure and the axial compressor for low pressure of the mixed refrigerant for International Scientific Electric Power Conference \u2013 2019 IOP Conf. Series: Materials Science and Engineering 643 (2019) 012149 IOP Publishing doi:10.1088/1757-899X/643/1/012149 example. Due to the large flexibility of speed range of the induction motor, it is possible to use identical electric systems for all compression services. Thanks to the development of high-speed induction motors and active magnetic bearings, integrated moto-compressors (see fig.10) represent today an alternative solution to conventional compression for SSLNG. The process gas handled by the compressor is used to cool both the motor and the magnetic bearings making the unit fully hermetic, allowing to remove the gas dry seals and the associated conditioning system, and reducing the footprint by 40% [12-16]. In addition, the integrated motor compressors are not sensitive to the load start conditions and do not need to modify the processing conditions to lower the Settle Out Pressure (SOP) before starting" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002776_isef45929.2019.9096891-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002776_isef45929.2019.9096891-Figure3-1.png", + "caption": "Fig. 3. Stress and deformation analysis of rotor disk using FEA.", + "texts": [ + " After assembling the AFPM motor, the rotor disks are integrated with a shaft suspended by magnetic bearings both radial and axial directions. Total system is shown in Fig. 2. Analytical solutions are omitted due to limited space. In order to validate the analytical model results, FEA are performed. Mechanical properties of the materials used in FEA is tabulated in Table 2. III. RESULTS AND DISCUSSION In this section a detailed comparison between analytical model results and FEA will be added in full-paper; for digest paper, only the preliminary results are given in Table 3 and Fig. 3. Preliminary results show that the first natural mode is higher than rated operating speed which can be interpreted as shaft can be rotated safely. The results of the mechanical stress analysis (MSA) show that the bending values of outer edges of rotor disks are much lower than the critical limits. IV. CONCLUSION This paper presents the mechanical vibration and stress analysis of an AFPM motor with a high speed rotor that is vertically suspended by active and hybrid magnetic bearings. Modal analysis of the total system with shaft, thrust disk and rotor disks are performed using both analytically and FEA" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001629_012016-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001629_012016-Figure2-1.png", + "caption": "Figure 2. Illustration of composite stacking and their principal axes", + "texts": [ + " Series: Materials Science and Engineering 700 (2019) 012016 IOP Publishing doi:10.1088/1757-899X/700/1/012016 Lamina, or ply, is a plane (curved) layer of unidirectional fiber or woven fabric in a matrix. It is called unidirectional lamina in case of unidirectional fiber. Lamina is an orthotropic material with principal material axes in the direction of the fibers (longitudinal), normal to the fibers in the plane of the lamina (inplane transverse) and normal to the plane of the lamina. These principal axes are designated as 1, 2 and 3 respectively as shown in Figure 2. Laminate is a combination of two or more unidirectional laminae or plies stacked together at various orientations. The laminae can be of different thicknesses and consists of different materials, based on the orientation of the principal axes, laminates are analyzed using common coordinates (x, y, z). Lamina is a single flat layer of unidirectional fibers or woven fibers arranged in a matrix, whereas a laminate is a stack of plies of composites. Each layer can be laid at various orientations and can be made up of different material systems" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000236_0954407019838415-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000236_0954407019838415-Figure2-1.png", + "caption": "Figure 2. Finite element model of crankshaft.", + "texts": [ + " Crankshaft deformation and journal load The finite element method1,15 is used to solve the journal load of engine main bearing and the misaligned angle of crankshaft in bearing bore caused by the deformation of crankshaft in a stable working cycle of the engine, which based on the measured cylinder pressure and the inertial force of the crank-link mechanism, and the detailed solution process can be found by Sun et al.1 In order to improve the analysis quality and accuracy of the finite element of crankshaft, the mesh is refined in the vicinity of the shoulder, the transition fillet, and the oil inlet hole on the crankshaft journal. The finite element model of crankshaft is displayed in Figure 2, with a total of 266,283 nodes and 175,579 units. Figure 3 shows the No. 1 and No. 3 main bearing journal loads in a working cycle of the engine. Figure 4 shows the change of misaligned angle against the crankshaft rotate-angle when the crankshaft journal tilts in the bearing hole. Figure 5 shows the displacement of main bearing journal in the direction of bearing axis caused by the elastic deformation of crankshaft under different crankshaft angles. Measurement of axial movement of deformed crankshaft On the test bench of engine, the axial movement law of deformed crankshaft is measured by a displacement sensor installed at the free end of the crankshaft under the demarcated working condition of engine (3200 r/ min, 100% load), and the detailed test methods are shown in Sun et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001650_ijseims.2020010105-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001650_ijseims.2020010105-Figure6-1.png", + "caption": "Figure 6. Pad mesh", + "texts": [], + "surrounding_texts": [ + "=\ufeff3\ufeffmm\ufeffand\ufeffthe\ufeffdepth\ufeffof\ufeffry\ufeff=\ufeff0.5\ufeffmm,\ufeffthe\ufeffaxial\ufeffdistance\ufeffbetween\ufeffthe\ufefftextures\ufeffd\ufeff=\ufeff10\ufeffmm\ufeffand\ufefftheir\ufeff angular\ufeffoffsets\ufeff\u03b1\ufeff=\ufeff10\u00b0.\nMeshing The\ufefffinite\ufeffelement\ufeffnumerical\ufeffsimulation\ufeffis\ufeffused\ufeffto\ufeffcalculate\ufeffthe\ufeffdisplacement\ufeffof\ufeffthe\ufeffinner\ufeffface\ufeffof\ufeffthe\ufeff plain\ufeffbearing.\ufeffThe\ufeffsolid\ufeffis\ufeffdecomposed\ufeffinto\ufeffa\ufeffnumber\ufeffof\ufeff4-node\ufeffor\ufeff8-node\ufefftetrahedral\ufefffinite\ufeffelements\ufeff so\ufeffthat\ufeffthese\ufeffelements\ufeffare\ufeffas\ufeffaccurate\ufeffas\ufeffpossible\ufeffin\ufeffthe\ufeffgeometry.\nThe\ufeffshaft\ufeffis\ufeffdiscretized\ufeffinto\ufeffhexahedral\ufeffelements\ufeffwith\ufeff8\ufeffnodes,\ufeff(Figure\ufeff5),\ufeff15\ufeffnodes\ufeffin\ufeffthe\ufeffaxial\ufeff direction,\ufeff54\ufeffpoints\ufeffin\ufeffcircumferential\ufeffand\ufeff22\ufeffpoints\ufeffin\ufeffthe\ufeffradial\ufeffdirection.", + "The\ufeff pad\ufeff is\ufeff decomposed\ufeff into\ufeff 4\ufeff nodes\ufeff tetrahedral\ufeff elements\ufeff due\ufeff to\ufeff the\ufeff existence\ufeff of\ufeff a\ufeff groove\ufeffwith\ufefflines\ufeffand\ufeffarches\ufeffas\ufeffwell\ufeffas\ufefforifices\ufeffin\ufeffthe\ufeffcircle\ufeffform.\ufeffThese\ufeffshapes\ufeffrequire\ufefftetra\ufeff elements\ufeffto\ufeffachieve\ufeffthe\ufeffmost\ufeffaccurate\ufeffgeometry\ufeffoverlap,\ufeffFigure\ufeff6.\ufeffThe\ufeffbad\ufeffis\ufeffdecomposed\ufeffat\ufeff 12\ufeffpoints\ufeffin\ufeffthe\ufeffaxial\ufeffdirection,\ufeffthe\ufeffangular\ufeffamplitude\ufeffhas\ufeff70\ufeffpoints\ufeffand\ufeff4\ufeffnodes\ufeffdepending\ufeff on\ufeff the\ufeff thickness.\ufeff The\ufeff feed\ufeff groove\ufeff is\ufeff discretized\ufeff at\ufeff 13\ufeff points\ufeff in\ufeff the\ufeff radial\ufeff direction\ufeff and\ufeff 5\ufeffnodes\ufeffalong\ufeff its\ufeffwidth.\ufeffThis\ufeff finite\ufeffelement\ufeffmesh\ufeffof\ufeff the\ufeffplain\ufeffbearing\ufeffconsists\ufeffof\ufeff19830\ufeff elements\ufeffand\ufeff47784\ufeffnodes.\nBoundary Conditions To\ufeffbetter\ufeffresolve\ufeffthe\ufeffequations\ufeffpresented\ufeffearlier\ufeffin\ufeffthis\ufeffarticle,\ufeffit\ufeffis\ufeffnecessary\ufeffto\ufeffapply\ufeffappropriate\ufeff boundary\ufeffconditions.\ufeffThe\ufeffboundary\ufeffconditions\ufeffapplied\ufeffto\ufeffthe\ufeffshaft\ufeffare:\n\u2022\ufeff The\ufeffnodes\ufeffof\ufeffthe\ufeffface\ufeffwhere\ufeffthe\ufeffshaft\ufeffis\ufeffcoupled\ufeffwith\ufeffa\ufeffrotating\ufeffmanifold\ufeffare\ufeffblocked\ufeffalong\ufeffthe\ufeff x-axis,\ufeffFigure\ufeff7.\nThe\ufeffboundary\ufeffconditions\ufeffapplied\ufeffto\ufeffthe\ufeffpad\ufeffare:\n\u2022\ufeff The\ufeffpad\ufeffis\ufeffplaced\ufeffin\ufeffa\ufeffsupport\ufeffring\ufeffwhose\ufefflower\ufeffpart\ufeffis\ufeffspherical,\ufeffa\ufeffframe\ufeffone\ufeffbears\ufeffon\ufeffthe\ufeffface\ufeff of\ufeffthe\ufeffspherical\ufeffhydrostatic\ufeffplain\ufeffbearing\ufeffand\ufeffwhich\ufeffis\ufeffembedded\ufeffin\ufeffthe\ufeffbase,\ufeffFigure\ufeff8; \u2022\ufeff The\ufeffpad\ufeffis\ufefflocked\ufeffon\ufeff60\u00b0\ufeffof\ufeffthe\ufefflower\ufeffpart:\ufeffBlocking\ufeffof\ufeffthe\ufeffnodes\ufeffaccording\ufeffto\ufeffx,\ufeffy\ufeffand\ufeffz; \u2022\ufeff The\ufeffpad\ufeffis\ufeffmounted\ufeffin\ufeffa\ufeffring\ufeffso\ufeffit\ufeffis\ufeffcylindrical\ufeffsupport\ufeffwith:\n\ufeff\u25e6 Fixed\ufeffradial\ufeffnodes; \ufeff\u25e6 Free\ufeffAxial\ufeffNodes; \ufeff\u25e6 Fixed\ufefftangential\ufeffnodes.\nInsertion of Pressures We\ufeffcan\ufeffapply\ufeffglobal\ufeffloads,\ufeffstructural,\ufeffas\ufeffwell\ufeffas\ufeffimposed\ufeffdisplacements\ufeffaccording\ufeffto\ufeffthe\ufeffcases\ufeffstudied.\ufeff In\ufeffthe\ufeffcase\ufeffstudied,\ufeffpressures\ufeffare\ufeffapplied\ufeffalong\ufeffthe\ufeffcircumferential\ufeffaxis\ufeffas\ufeffwell\ufeffas\ufeffalong\ufeffthe\ufeffaxial\ufeffaxis\ufeff of\ufeffthe\ufeffshaft\ufeffand\ufeffthe\ufeffplain\ufeffbearing\ufeff(Figure\ufeff9).", + "Radial Load Effect Pressure Distribution for Textured Plain Bearing To\ufeff demonstrate\ufeff the\ufeff effect\ufeff of\ufeff the\ufeff radial\ufeff load\ufeff on\ufeff the\ufeff operating\ufeff performance\ufeff of\ufeff the\ufeff non-textured\ufeff hydrodynamic\ufeffplain\ufeffbearing,\ufeffsuch\ufeffas\ufeffthe\ufeffpressure\ufeffdistribution\ufeffand\ufeffthe\ufefffluid\ufeffflow\ufeffvelocity\ufeffwithin\ufeffthe\ufeff plain\ufeffbearing,\ufeffa\ufeffradial\ufeffload\ufeffvariation\ufeffis\ufeffperformed\ufeff(W1\ufeff=\ufeff2000N,\ufeffW2\ufeff=\ufeff6000N\ufeffand\ufeffW3\ufeff=\ufeff10000N).\ufeff The\ufeffinitial\ufeffoperating\ufeffconditions\ufeffof\ufeffthe\ufeffbearing\ufeffare\ufeffwith\ufeffa\ufeffsupply\ufefftemperature\ufeffTa\ufeff=\ufeff40\ufeff\u00b0C,\ufeffsupply\ufeff pressure\ufeffPa\ufeff=\ufeff0.04\ufeffMPa\ufeffand\ufeffthe\ufeffspeed\ufeffof\ufeffrotation\ufeffof\ufeffthe\ufeffshaft\ufeffequal\ufeffto\ufeffN\ufeff=\ufeff6000\ufeffrpm." + ] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure82.7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure82.7-1.png", + "caption": "Fig. 82.7 CA GCI and AlSiC profile of von Mises stress (Model 1)", + "texts": [], + "surrounding_texts": [ + "In the coupled analysis, the thermal load was coupled with the structural load to find out the combined effect on brake disc models. The temperature induced at various time points was imported into static structural, and then structural loads were applied. The analysis was run for 36 s, i.e. the time taken by the vehicle to stop due to the application of the emergency brake. The output results of von Mises stress and total deformation developed in the model were recorded (Figs. 82.7, 82.8 and 82.9). Some important points that can be drawn from the analysis are: 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 979 \u2022 When compared to the Factor of Safety offered by the GCI models, the AlSiC models offer higher Factor of Safety. \u2022 For the same applied load, the AlSiC models have lower thermal stresses than the GCI models, as AlSiC material has greater thermal conductivity and heat dissipation capability. \u2022 The weight of AlSiC models is lesser when compared to the GCI models (GCI model having a weight of about 134 kg gets reduced to 54 kg in case of AlSiC material). 980 E. Madhusudhan Raju et al." + ] + }, + { + "image_filename": "designv11_80_0001288_j.matpr.2019.08.229-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001288_j.matpr.2019.08.229-Figure3-1.png", + "caption": "Fig. 3. Internal pressure distribu", + "texts": [ + " Thin cylinders design mainly involves the calculation of thickness (t) as Cylindrical cylinder thickness; t \u00bc pd 2rc \u00f03:3\u00de tion of cylindrical cylinder. (b) Sandwich reinforced composite. and FE analysis of epoxy composite pressure cylinder used for aerospace 9.08.229 Fig. 4. (a) 2-D Model of pressure cylinder, and (b) Model developed in CREO parametric. Fig. 5. (a) Pressure cylinder model, (b) Bonded region, (c) Fixed support, and (d) Internal pressure. Internal pressure distribution of cylindrical cylinder is shown in Fig. 3. Circumferential stresses will be rc \u00bc Total pressure Resisting section \u00bc pdl 2tl \u00bc pd 2t \u00f03:4\u00de The longitudinal stresses will be, rt \u00bc Total pressure Resisting section \u00bc pd 4t \u00f03:5\u00de Now changes in diameter and length may be found out from the above equations, dd \u00bc TM 1 d \u00bc pd 2tE 1 1 2m d \u00f03:6\u00de dl \u00bc TM 2 l \u00bc pdi 2tE 1 2 1 m \u00f03:7\u00de Please cite this article as: K. Rajendra Prasad and C. Syamsundar, Theoretical and FE analysis of epoxy composite pressure cylinder used for aerospace applications, Materials Today: Proceedings, https://doi" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001205_012017-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001205_012017-Figure1-1.png", + "caption": "Figure 1. Positive deflection of control fins viewed from rear", + "texts": [ + " These fin deflections then act as a forcing function to the airframe dynamic model. Therefore, the main task of the autopilot subsystem can be summarized as follows: \u2022 Ensure the desired acceleration characteristics for tracking guidance commands with high performance with sufficient system stability. \u2022 Disturbance rejection in roll and pitch channels for the working flight envelop. The missile actuation can be achieved by accurately controlling the fins surfaces angles where the control commands require the convention of positive surface deflection angles as shown in Fig.1 because there are four fins (\u03b41, \u03b42, \u03b43, \u03b44), but only three attitude degrees of freedom (\u03be, \u03b7, \u03c2), they are combined, mathematically to form the three control commands as follows [3]: 1 2 3 4 1 2 3 4 1 2 3 4 1 ( ) 4 1 ( ) 4 1 ( ) 4 \u03be \u03b4 \u03b4 \u03b4 \u03b4 \u03b7 \u03b4 \u03b4 \u03b4 \u03b4 \u03c2 \u03b4 \u03b4 \u03b4 \u03b4 = + + + = \u2212 \u2212 + + = \u2212 \u2212 + (1) Four positive fin deflections create a negative roll command; the first two fins up (negative) and the last two fins up (positive) generate a positive normal force command; and the fins 2&3 negative deflections with fins 1&4 positive deflections cause a positive yaw command", + " the trim points should be at different flight conditions during the rocket motor powered and unpowered stages to avoid repeating the design points [3]. In order to select the design points, the Mach number and altitude must be plotted with the flight time, or instead of them the dynamic pressure can be introduced with the change of flight time and trajectory as shown in Fig. 2. 18th International Conference on Aerospace Sciences & Aviation Technology IOP Conf. Series: Materials Science and Engineering 610 (2019) 012017 IOP Publishing doi:10.1088/1757-899X/610/1/012017 As shown in Fig.1: \u2022 At the first seconds the change in altitude and dynamic pressure is almost small. \u2022 The dynamic pressure reaches its maximum value at the end of powered phase at time (t=13.05 sec). \u2022 The dynamic pressure decreases at the beginning of the unpowered phase. \u2022 Before the altitude reaches summit and at time (t=60sec), Qbar remains almost with constant value till time (t=120 sec) during passing through the summit. \u2022 After 120 sec the dynamic pressure returns to increase due to increasing of atmospheric density and velocity till reaching the target point" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000635_978-3-030-20377-1_1-Figure1.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000635_978-3-030-20377-1_1-Figure1.5-1.png", + "caption": "Fig. 1.5 Surface displacements due to surface forces", + "texts": [ + " We specialise the results to the surface with the outcome u(x) = \u2212 P [ (1 \u2212 2\u03bd) (1 + \u03bd) 2E ] sgn(x)+ + Q [ 2 (1 \u2212 \u03bd) (1 + \u03bd) \u03c0E ] ln |x | + c1 (1 + \u03bd) E (1.9) v(x) = \u2212 P [ 2 (1 \u2212 \u03bd) (1 + \u03bd) \u03c0E ] ln |x |+ + Q [ (1 \u2212 2\u03bd) (1 + \u03bd) 2E ] sgn (x) + c2 (1 + \u03bd) E , (1.10) where v(x) is the surface normal displacement, u(x) is the surface tangential displacement, E is Young\u2019s modulus, and \u03bd is Poisson\u2019s ratio. These results merit comment; first, observe that, as we might intuitively expect, the normal force, P , causes a logarithmically varying surface normal depression, Fig. 1.5, with the depth of the depression becoming unbounded as the point of application of the load is approached. Perhaps rather less obvious is that the normal force also causes surface particles to be drawn in, parallel with the surface, towards the loaded point, and with the magnitude of the displacement independent of the position of the particle being considered. Similarly, the shear force causes surface particles to be pushed/drawn along with a magnitude varying logarithmically with position. But note, also, that ahead of the shear force material is pushed down by a constant amount whilst behind it it is raised, forming a step, Fig. 1.5. Lastly, note that each equation incorporates an arbitrary constant so that some datum depth must be chosen at which the displacements are arbitrarily set to zero. Partly to eliminate the need for an arbitrary datum depth, we develop a contact formulation in terms of the gradient. So, it is straightforward to find the surface slope, dv/dx , at point x due to an element of pressure, p(\u03be)d\u03be at point \u03be , but the effect of the shear force needs a little more thought. If the shear force, Q, is smeared over a small length dx (and Q = q(x)dx), we see that the surface slope is affected only by the surface shear traction at that point, and not elsewhere; by contrast the contact pressure at any point influences the slope at all points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002124_slct.201904873-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002124_slct.201904873-Figure3-1.png", + "caption": "Figure 3. Schematic presentation of Cu(acac2pn)/GOx/CPE biosensor.", + "texts": [ + " We presume that the hydrophobic/hydrophilic properties of this Schiff base Cu(II) complex enable the penetration of this mediator close to the enzyme redox site. Finally, the reduced mediator is electrochemically re-oxidized on the electrode surface, causing the change in the oxidation current. The catalytic process involves the enzymatic transformation of the glucose and the mediator. Every electron transferred to the surface of electrode corresponds to one molecule of glucose being oxidized. The schematic illustration of functioning of this mediate biosensor is shown on Figure 3. Electrochemical behavior of Cu(acac2pn) were determined by cyclic voltammetry in DMSO/PBS pH 6.6 solution. From the obtained voltammogram (Figure 4) the peak at 0.0 V can be attributed to the copper oxidation and corresponding copper reduction peaks ( 0.175 and 0.25 V) which were more pronounced after the addition of copper-acetate.[28,34] 1672ChemistrySelect 2020, 5, 1671\u20131675 \u00a9 2020 Wiley-VCH Verlag GmbH & Co. KGaA, Weinheim Wiley VCH Montag, 03.02.2020 2005 / 157590 [S. 1672/1675] 1 In mediated electron transfer (MET), electrons transfer from enzyme active site to the electrode surface by mediator who acts as an electron relay" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001051_ccdc.2019.8832344-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001051_ccdc.2019.8832344-Figure2-1.png", + "caption": "Fig 2.The axes system used in this paper.", + "texts": [ + " Finally, a simulation is designed to verify the whole controller, whose results reach the request. Key Words: Tilt-Wing UAV, Robust Attitude Controller, Optimal Control Allocation, DOB 1480978-1-7281-0106-4/19/$31.00 c\u00a92019 IEEE Authorized licensed use limited to: University of Exeter. Downloaded on May 05,2020 at 19:31:29 UTC from IEEE Xplore. Restrictions apply. 2.2 The Axes System The axes systems in this paper include the body axes system b b bOx y z and the Earth-fixed axes system g g gOx y z , that is, the North-East-Down (NED) system. Fig 2 shows the directions and relationship of these axes systems. 2.3 The Dynamic Model A nonlinear 6 degree-of-freedom rigid-body dynamic model is established based on the Newton\u2019s law of motion: ( )B B B BF m v v\u03c9= + \u00d7 (1) ( )B B B BM J J\u03c9 \u03c9 \u03c9\u00d7= + (2) I B vv R v= (3) I B wR\u03c9 \u03c9= (4) where [ , , ]B Tv u v w= , , ,[ ]I Tv x y z= , [ , , ]B Tp q r\u03c9 = , , ,[ ]I T\u03c9 \u03c6 \u03b8 \u03c8= denote the velocity and the angular velocity vectors in body axes and NED axes system. , ,v wJ R R denote the inertia rotation matrix for velocity and angular velocity between body axes and NED axes system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002043_iros40897.2019.8968006-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002043_iros40897.2019.8968006-Figure2-1.png", + "caption": "Fig. 2. Model of the femoral component including the stimulation electrode and insulator.", + "texts": [ + " The hemispherical stimulation electrodes have a diameter of 4 mm, while the ovoid cup has a length of 70 mm and a width of 58 mm. The acetabular cup also includes an anchorage cone, which is placed in one of five default drill holes within the cup. The femoral component used in this study is a preliminary prototype still in its design state. So far, the component is composed out of a hip stem (Hipstar, size 2)1 with a notch on one side of the implant, in which the stimulation electrode (width: 1.45 mm) and the surrounding biocompatible insulator are located (see Fig. 2). 1http://bizwan.com/_mydoc/stryker/Hip/HipStar TMZF Cementless Hip System Surgical Technique.pdf. A primary coil, which is placed around the patient\u2019s hip, provides a time-harmonic magnetic field at a frequency of 20 Hz. This oscillating field induces a location-dependent current in the secondary coils of the stimulation electrodes of the acetabular cup and the femoral component resulting in an electric potential distribution in the area around the implants. During surgery, each electrode of the acetabular cup has to be placed into the bone, which limits the number of utilizable stimulation electrodes", + " The accelerating effect on the growth of bone cells by electrostimulation described by Basset et al. [3] depends on the electric field distribution in the bone. This field distribution is determined by the placement of the stimulation electrodes on the implant and depends also on the geometry of the pelvic bone as well as the femoral part and the individual defects as classified in [14]. While the femoral component of the implant has a fixed stimulation electrode located in a notch of the implant (see Fig. 2), the stimulation electrodes on the acetabular cup can be placed at different locations to provide a preferably large beneficial stimulation region. To provide the optimal electrode arrangement for the acetabular cup, we apply a multiobjective evolutionary optimization algorithm as presented in [6]. Based on the work of Kraus [16] and clinical advice, an electric field of 5\u201370V/m in the proximity of the implant and 35\u201370V/m in the defective area, which both forms the area of interest, is considered to be optimal" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001090_032034-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001090_032034-Figure4-1.png", + "caption": "Figure 4. The minimum allowable values of indicators of the relative position of the total contact patch L1, L2, L3, L4.", + "texts": [ + " 2), as a result of which it is unsuitable for developing comprehensive recommendations for quality improvement [6] Therefore, before developing a method for controlling the quality of the manufacture of gear rims in the toothing process, a method was developed to assess the compatibility of the gear teeth of the differential satellite in terms of both dimensions and the relative position of the total contact patch. As a result, in addition to the standard indicators characterizing the size of the contact patch relative to the length and height of the tooth% H,% L (fig. 3), the indicators L1, L2, L3, L4 were added - the distances from the extreme points of the total contact patch to the borders of the active tooth surface (fig. 4, 5). Mechanical Science and Technology Update IOP Conf. Series: Journal of Physics: Conf. Series 1260 (2019) 032034 IOP Publishing doi:10.1088/1742-6596/1260/3/032034 To find the whole complex of operating factors affecting the magnitude of the total contact patch, taking into account the results of a positive solution of similar problems in various processes of forming machine parts [7\u201311], the main key quality indicators were found, the values of which should be measured when production experiment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000505_14484846.2019.1626529-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000505_14484846.2019.1626529-Figure13-1.png", + "caption": "Figure 13. The pressure contour maps in the gas films of orifice aerostatic bearings.", + "texts": [ + " The restrictors parameters are based on the results of the previous optimization. From the above analysis, it can be seen that the pressure peak appears on the thrust surface of the orifice-type bearing. In order to make the force more uniform, the equalizing-pressure groove with the section of equilateral triangle is set on the thrust surface, and the depth of the groove is 0.15 mm. Similarly, the same boundary conditions as above are set to optimize the number of restrictors. The pressure contour maps in the films can be obtained by CFD. It can be seen from Figure 13 that the highpressure area within the gas film is concentrated near orifices and the equalizing-pressure groove, and the pressure falls to atmospheric pressure at the bearing edge. With the increase of the number of restrictors, the area of high pressure in the gas film increases. Figure 14 shows the static characteristic curves of bearings with a different number of orifices. The following conclusions can be drawn from the analysis: (1) Increasing the number of orifices can improve the load capacity of the bearing, which cannot be greatly improved if the number of orifices is greater than 8" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003785_etfa46521.2020.9212115-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003785_etfa46521.2020.9212115-Figure2-1.png", + "caption": "Fig. 2: The triangle formed. L is the source, S is the sensor and P is the projection.", + "texts": [ + " 1) Basic Scheme: The main idea is to use the RSSI measurements of four sensors in order to calculate the position of the electromagnetic source in the 3d space. These sensors (S0, SR, SL, SF ) should be placed on the perimeter of a circle with radius r as seen on Fig. 1. Each sensor returns a measurement E, defined by (1), which is proportional to the intensity of the source (W ) and the angle of incidence (\u03c6 ) and inversely proportional to the square of the distance between the sensor and the source. E = W 4\u03c0d2 cos\u03c6 (1) The triangle formed by the sensor, the source and the projection of the source onto the sensors plane can be seen on Fig. 2. The cosine of the angle of incidence is equal to the 683 Authorized licensed use limited to: Carleton University. Downloaded on November 03,2020 at 11:55:26 UTC from IEEE Xplore. Restrictions apply. sine of the angle formed by d and b (cos\u03c6 = sin (90\u2212 \u03c6)). Thus equation (1) can be written as: E = W 4\u03c0d2 sin (90\u2212 \u03c6)\u21d2 E = W 4\u03c0d2 h d \u21d2 E = Wh 4\u03c0d3 (2) Solving equation (2) for d, we come up with equation (3) where \u03b2 = Wh 4\u03c0 . d = \u03b2 1 3 E 1 3 (3) Working on the robot-centric system, the coordinates of each sensor are as seen on Table I, where (xs, ys, zs) is the position of the source" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003739_s12239-020-0106-8-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003739_s12239-020-0106-8-Figure7-1.png", + "caption": "Figure 7. Schematic diagram about mechanical analysis of wheel external structure under braking condition.", + "texts": [ + " Mechanical Model of Modular Deformable Tire Under Braking Condition The mechanical behavior of modular deformable tires under braking conditions is similar to that of driving conditions. When braking torque is applied, the wheel still keeps moving forward, while the rubber block in contact with the ground produces a large elastic deformation, and the rolling resistance moment can be ignored (Zhisheng, 2006). Assume the acceleration and angular acceleration of the tire under braking are respectively a' and \u03b1'. Figure 7 is a schematic diagram of tire mechanics analysis under braking condition. Q1 Q2 P1 1 2 1 1 1 Q Q S P T P X i X1 6 1 i i 2 X P Pi 1 2 1 11 12 6 1 6 1 6 Q Q S P Q Q S P T P P X X i i Q11 Q12 X 6 5 1 6 i i 2 X X P (6) where Mb represents the braking moment; fb represents the braking force provided by the ground to the tire; represents the horizontal force acting on the tire by the axle under the braking condition. Under braking condition, the supporting force T provided by the ground to the tire is transferred to the petalshaped connecting block, and the pressure generated by the piston fixed to the petal-shaped connecting block on the air cavity is the same as that under driving condition, so force analysis is not carried out here" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003695_s11041-020-00566-5-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003695_s11041-020-00566-5-Figure1-1.png", + "caption": "Fig. 1. Schematic diagram of the electron-beam drip melting (EOS \u2014 electron-optical system): 1 ) horizontal feeding rod of the source metal; 2 ) generated drops; 3 ) bath of melted metal in a water-cooled crucible.", + "texts": [ + " The aim of the present work is to determine the parameters of EBM promoting the removal of admixtures in the process of refinement of various chemically active and nonferrous metals and in the course of reprocessing of the scrap of Cu, Hf, and Zr and to perform the analysis of the results of application of the EBM method. Metal Science and Heat Treatment, Vol. 62, Nos. 5 \u2013 6, September, 2020 (Russian Original Nos. 5 \u2013 6, May \u2013 June, 2020) 345 0026-0673/20/0506-0345 \u00a9 2020 Springer Science+Business Media, LLC 1 Institute of Electronics, Bulgarian Academy of Sciences, Sofia, Bulgaria. 2 E-mail: katia@van-computers.com. DOI 10.1007/s11041-020-00566-5 Specific features of the process of electron-beam melting. In Fig. 1, we present a schematic diagram of the EBM (electron beam drip melting) method. The process of melting is carried out in a vacuum chamber. The formed electron beam is directed by one (or several) electron-optical system into the reaction zones: to the remelted rod 1 and to the surface of liquid bath 3 of the ingot. The electrons collide with the refined metal and heat it. The drops of melted metal 2 fall into the water-cooled copper crucible with moving bottom, where they are solidified. The surface of melted metal in the crucible is also heated by the electrons. The processes of re- finement mainly occur on the melted-metal vacuum interface in the front part of the source (raw) material. The feeding rod creates drops and the liquid bath in the crucible (Fig. 1). The main parameters of the process controlled by the operator are as follows: startup power, focusing current of the beam, melting rate and or casting rate (or the rates of feeding and drawing), as well as the cross section and density of the supplied metal. By varying the power of the beam and casting rate, it is possible to control the geometry of the bath of liquid metal, the distribution of temperature in the zone of interaction of the beam with the metal, and the duration of the thermal influence affecting the processes of refinement and the quality of the metal obtained as a result of EBM. The choice of the corresponding parameters of the technological processes guarantees the possibility of creation of the thermodynamic, hydrodynamic, and kinetic conditions required for the removal of various admixtures. In the present work, all experiments were carried out with the use of electron beams with a power of 60 kW (ELIT 60 installation) with a single electron gun, a horizontal feeder, a water-cooled copper crucible, and a withdrawal system (pulling mechanism) (Fig. 1). Application of EBM as the final stage of an ordinary metallurgical processes of production of pure metals. The indicated possibility of application of the EBM method makes it possible to exclude numerous intermediate technological operations some of which are hazardous for the environment and dangerous for health. In many cases, it is possible to include EBM in the earlier stages of the ordinary metallurgical process. In this case, metals with high contents of admixtures can be used as source (raw) materials" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000019_3305275.3305335-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000019_3305275.3305335-Figure5-1.png", + "caption": "Figure 5. Spherical joint: (a) Model of spherical joint; (b) Object of spherical joint", + "texts": [], + "surrounding_texts": [ + "2.1 The Composition and the Finite Element Model Parameters The ADAM consists of a series of identical span units, each of which is composed of transverse bar, longitudinal bar, spherical joint, stay cable component and guide wheel, as shown in Figure 2. Transverse bar and longitudinal bar are connected through spherical joint, and guide wheel ensures the ADAM to expand smoothly, while stay cable component provides tension force for the mast. According to actual application, the geometrical dimensions of transverse bar and longitudinal bar are shown in Table.1. ANSYS a high-efficiency finite element analysis software integrating structure, fluid, heat, magnetic, etc., and applicable to the ADAM described in this paper, which can improve the efficiency of simulation greatly. In the ANSYS environment, physical properties and material properties of each component unit are set as shown in Table 2. Further, for the stay cable, the linear density is 10.33g/m, so the equivalent cross-sectional area is 1.3076\u00d710-6m2, while the locking device and the spherical joint are set as mass points of 0.1899kg and 0.150kg respectively. The material of transverse bar and longitudinal bar is special material for carbon fiber/epoxy resin composites with the grade M55; the stay cable is made of stainless steel material of the grade 1Cr18Ni9Ti. 2.2 Process of Parameterized Finite Element Analysis of ADAM The parametric analysis process is as follows: firstly, deterministic parameters of ADAM are obtained to establish mast model in ANSYS quickly, and the unit type, material properties and physical properties of each part are set secondly. Then constraint conditions are added on the meshed model, such as fixing mast, setting joint stiffness, enforce pre-tension, etc. After completing simulation calculation, results finally are viewed and post-processed. The simulation process is shown in Figure 3. Based on physical performance parameters of mast, such as span number, material, pre-tension of stay cable stiffness of joint and so on, the secondary development of ANSYS is used to complete modal analysis of mast by using VB6.0 in this paper. As shown in Figure 4, the secondary platform software can automatically call ANSYS program to complete mesh division, constraint addition, and solution by entering accurate parameters according to the user interface of the software." + ] + }, + { + "image_filename": "designv11_80_0003316_icced46541.2019.9161127-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003316_icced46541.2019.9161127-Figure8-1.png", + "caption": "Fig 8. Wiring the switch with the relay in series", + "texts": [ + " Where users be able monitor electrical power usage, and be able control and monitor the use of home appliances that consume electrical power through a relay that functions as a switch that can be controlled remotely. How to control the relay through the button as the virtual switch that has been provided on the functioned website is worth 1 or 0. A value of 1 on the button indicates the relay is on, so if the button is 0 then the relay is turned off. Authorized licensed use limited to: Cornell University Library. Downloaded on August 28,2020 at 13:55:20 UTC from IEEE Xplore. Restrictions apply. Installation of a relay with a analog switch is arranged in series as shown in the following Figure 8. In Figure 8, in using the switch manually, the relay must be Normaly Close (NC) or wiring in the NC slot position and the virtual switch represented by the button is 0. Similarly, when using the virtual switch control, the manual switch must be connected or position 1 or close. If the manual switch is in the 0 or open state, then the use of a virtual switch that controls the relay cannot work or cannot process ON or run electricity so that electrical equipment cannot be turned on remotely. Because the results of this study are in the form of prototypes, it is necessary to treat them for testing purposes" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001739_ismsit.2019.8932756-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001739_ismsit.2019.8932756-Figure5-1.png", + "caption": "Fig. 5. Isometric View of Final Design", + "texts": [], + "surrounding_texts": [ + "III. SOCIO-ECONOMIC SIGNIFICANCE\nAt present farming industry make use of drones for surveillance and monitoring fields these drones are equipped with spectrum sensors that can identify the crop type and its condition. Then there are the conventional farming methods that employ manual labor. Also aerial sprays are used to spray fields with fertilizer and pesticides. Similar Machinery exists but is only household specific; however we are aiming to target research and development departments in pharmaceutical and food industry, where setting up small greenhouse is required. Further we would be using some tool head for multiple operations instead of relying on magnetism for tool selection. Cutting down cost is another objective by using alternative materials to steel.\nIV. DESIGN\nWhile developing the project we went through some major and minor issues which lead us to make multiple iterations in the design phase. In this section we will discuss these iterations.\n Iteration 1:\nThis is the initial idea and concept on which we started to work. Our initial idea was to have x-axis, y-axis, z-axis and one rotary axis for the axis of motion. We also thought about adding multi-head mounter and add moisture sensor, shower and seeder. Initially we want to use rack and pinion mechanism. Issues: Rack and pinion mechanism can increase the vibrations which can affect the structure.\nWe then start to work on to select a mechanism for x-axis. We start to work on the stepper motor and lead screw mechanism. Issues: Our project has an open base so lead screw can\u2019t be mounted at the bottom like every CNC milling machine and usage of two stepper motor for just one axis can effect or destroy the structure if something happens to one motor.\n Iteration 3: So we started to think more about it and started to work on another solution to move the x-axis with just one motor. We finally decide to move the whole axis with one motor by using belt and pulley mechanism as shown in figure below.\nIn this project we use t-slotted extruded bars for developing our structure. Y-axis consist of simple lead screw mechanism.\nZ-axis consist of simple lead screw mechanism for linear motion. On z-axis a servo motor is mounted for giving a rotary motion to the multi head. The final design after the iterations is shown below.", + " Vacuum Pump: When a signal is sent out from controller to the vacuum pump and vacuum pump turn on and suck air and move towards seed when it come near the seed, then seed float and got stuck on the nozzle until the vacuum is on. When it reaches to the desired location then the controller sends another signal and vacuum pump turns off and the seed drops on the location the grid.\n Water Pump: When a signal is sent out from the controller to the water pump then water pump turn on and sprinkle the water over the grid with the help of shower for some specific time and then another signal comes from the controller and turns it off. Piping and instrumentation diagram (P&ID) is shown below:\nCurrently there are lots of global challenges which we are facing like global warming, food needs, poverty, degradation of climate and much more. Sustainable development goals introduced by United Nation helps us address these problems and base on these problems find solution to make future better not just for humans but also for the creature that exist on globe. The goals which can be achieved by our project \u201cGarden Tech\u201d are as follows:\n Zero Hunger: As we all know that human population is increasing rapidly to accommodate them deforestation is taking place. It is effecting our climate and results in global warming due to which glaciers are melting. Flooding is taking place more often than ever before and leads to devastation of arable land. Due to these circumstances food growth is not increasing rapidly like human population. The anticipated population according to United Nations Department of Economic and Social Affairs (UN DESA) is shown below:", + "As shown in figure below lots of farm produces are destroyed due to drought, flood, pest attack and etc. Garden Tech is one of the best solution to overcome these sort of problems. Garden Tech can water the plants on timely manners as well as depends on moisture level so, it can produce quality food.\nGarden Tech can be implemented on the rooftop gardens, backyard, terraces and many other places so, every house can produce their own food which will be completely hygienic, healthy and free from harmful pesticides. If majority of the population started to do gardening in their homes than it will decrease the price of food and the food will be available to everyone which will lead to zero hunger.\n Good Health and Well-Being: As we know that health is wealth and good heath can lead us to a good and longer life and bad health can be a cause of death. There are lots of causes for human deaths like malaria, AIDS and many other. Hygiene and pollution problem is also a bigger cause for the health issues. Now a days foods are not that much hygienic and are growing with the help of pesticides which can grow them rapidly but are very harmful for human health. Plants and trees are one of the biggest source to clean air pollution but deforestation is the main cause that our climate is changing more often and global warming is taking place as discussed earlier. A need of quality medicines is also arises due to these conditions and a lot of medicines are made from herbs like\nherbal medicines. These medicines can only work if the herbs and plants are of good quality, hygienic and healthy. Now a day\u2019s people have lots of tensions which are leading them to diseases. Greenery is one of the best solution for humans to relax their mind but quantity of gardens, parks and jungles are decreasing. Each human being have a tough routine and don\u2019t have time to go to park which is far away from home. These causes are leading human to more stress which are the basic cause of headache, angry outburst, restlessness, chest pain, irritability, sleep problem and sadness or depression which can be a cause of suicide.\nGarden Tech is one of the best solution for this problem like Garden Tech can help in pharmaceutical companies in the development of good quality herbal medicines, research and development departments to work on plants hybrids, rooftop gardens, backyard or terrace gardens can help the getting fresh morning air and oxygen which is beneficial for human mind and can lead to happy life. It can reduce global warming and can reduce stress.\n Sustainable Cities and Communities: Human lands are divided into two rural and urban lands. Rapid urbanization management is now became an acute challenge for the cities around the globe. A need arises to support this increasing population which requires shelter, cloth and food as a basic requirement. Due to this huge amount of population, environment pollution is also increasing day by day which our leading our world to more disasters." + ] + }, + { + "image_filename": "designv11_80_0002015_s00419-020-01661-y-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002015_s00419-020-01661-y-Figure4-1.png", + "caption": "Fig. 4 Time history of active distortions exerted on the worm with 2p-period. a The relation between \u03b51(t) and t in the interval [0, L0]. b The relation between \u03b52(t) and t in the interval [L0, L]", + "texts": [ + " (5) and (7), one has T \u2032(Z , t) = \u23a7\u23a8 \u23a9 \u2212\u03bc\u2212, if u\u0307 < 0, \u2212 fv, if u\u0307 = 0, \u03bc+, if u\u0307 > 0, (8) where fv \u2208 [\u2212\u03bc\u2212, \u03bc+ ] and the tension expression can be written by T (Z , t) = K (u\u2032(Z , t) \u2212 \u03b50(t)). (9) Then, the positional relation between any two points Z1 and Z2 on the worm body is derived as u(Z2, t) = u(Z1, t) + \u222b Z2 Z1 (T (Z , t)/K + \u03b50(t))dZ . (10) If u\u0307(Z , t) = 0 for \u2200Z \u2208 U0(Z0), where U0(Z0) denotes a neighborhood of Z0, then it follows from Eq. (9) that T\u0307 (Z0, t) = \u2212K \u03b5\u03070(t). (11) Meantime, consider that the time history of active distortions is given by the 2p-periodic quadratic parabola graph in Fig. 4, namely the acceleration of its change is piecewise constant. Here, the active distortion in positional interval [0, L0] is typically given as \u03b50(t) = \u03b51(t). The other one [L0, L] can be denoted as \u03b50(t) = \u03b52(t). Then one obtains \u03b51(t) = \u03b11(t 2 \u2212 2pt), t \u2208 [0, 2p], (12) and \u03b52(t) = \u2212\u03b12(t 2 \u2212 2pt), t \u2208 [0, 2p]. (13) Actually, \u03b50(t) can indicate various active distortions (e.g., sawtooth form [18], sinusoidal form [1]). To deduce the net displacement of worm motion over one period based on Eqs. (8) and (9), the process from the start to the end of the motion over one period is specifically given as depicted in Fig. 4. Here, the positions of several special points are listed as follows XL1 = L0 + \u03bc+ \u03bc+ + \u03bc\u2212 (L \u2212 L0), XR1 = \u03bc\u2212 \u03bc+ + \u03bc\u2212 L0, XL2 = \u03bc+ \u03bc+ + \u03bc\u2212 L0, XR2 = L0 + \u03bc\u2212 \u03bc+ + \u03bc\u2212 (L-L0), (14) where XL2 + XR1 = L0, XL1 + XR2 = L \u2212 L0. (15) In fact, these points can be derived when the tension reaches the maximum value. Then it is seen from Fig. 5 that the symbol of the velocity u\u0307(Z , t) can be assigned. Namely, the interval [0, L] is partitioned into six disjoint sub-intervals wrote by [0, L] = IL2(t) \u222a I01(t) \u222a IR1(t) \u222a IL1(t) \u222a I02(t) \u222a IR2(t). (16) Here, these sub-intervals are presented according to the following two cases Case 1 Worm moves forward when the body center of mass moves backward. Then one has \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 u\u0307(Z , t) > 0, for Z \u2208 IL2(t) u\u0307(Z , t) = 0, for Z \u2208 I01(t) u\u0307(Z , t) < 0, for Z \u2208 IR1(t) \u23ab\u23ac \u23ad if \u03b5\u03071(t) < 0, u\u0307(Z , t) < 0, for Z \u2208 IL1(t) u\u0307(Z , t) = 0, for Z \u2208 I02(t) u\u0307(Z , t) > 0, for Z \u2208 IR2(t) \u23ab\u23ac \u23ad if \u03b5\u03072(t) > 0. (17) This case is shown Fig. 4 when t \u2208 [0, p]. Specifically, the tension reaches the maximal allowable value when t = \u03c41. However, the conditions \u03b5\u03071(t) < 0 for Z \u2208 [0, L0] and \u03b5\u03072(t) > 0 for Z \u2208 [L0, L] are still satisfied at this time. Hence, this case will continue until \u03b5\u03071(t) = 0 and \u03b5\u03072(t) = 0, and then the relation t = p is derived. In addition, it is observed from Fig. 4 that the other time interval t \u2208 [2p, 3p] is included in this case as well. However, since the tension of the fully relaxed state of a worm is equal to zero when t = 0, i.e., T (Z , 0) = 0, On the other hand, it can be obtained that T (Z , 2p) = \u2212\u03bc\u2212Z due to the continuity of the change of the tension in motion. It is shown that the initial value of the tension for t \u2208 [0, p] is apparently different from the one of the tension for t \u2208 [2p, 3p]. Therefore, the periodic motion of a worm does not start from the initial instant t = 0 Case 2 Worm moves backward when the mass center moves forward" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001857_033036-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001857_033036-Figure7-1.png", + "caption": "Figure 7. Scheme of forces in threaded connection: \u0430) moment of forces when screwing, \u0431) moment of forces when unscrewing", + "texts": [ + " Either action of constant force T (Figure 6 option 1) or a plane inclination relative to the horizon (Figure 6 option 2) that virtually does not differ from the first option can cause force asymmetry. When transverse vibration of the plane, in both cases, the body during semi oscillation either comes off the plane or the pressing force against it decreases. As a result the force T can move a body along axis x in spite of the fact that the body remained motionless without parametrical oscillations. We will determine the forces preventing unscrewing of threaded connections (Figure 7). The moment of forces in thread we will determine considering a nut as a slider which rising on turns of thread as on the inclined plane. According to formulas [4], moment needed for threaded connection unscrewing: \ud835\udc47\ud835\udc62\ud835\udc5b\ud835\udc60 = \ud835\udc47\ud835\udc61 + \ud835\udc47\ud835\udc52, (1) where Tuns \u2013 moment needed for connection unscrewing; Tt \u2013 frictional moment in turns of thread; Te \u2013 frictional moment on the end face. \ud835\udc47\ud835\udc61\ud835\udc56\ud835\udc54 = \ud835\udc39 \ud835\udc512 2 [tan(\ud835\udf11 + \ud835\udf13) + \ud835\udc53\ud835\udc47 \ud835\udc37\ud835\udc5a \ud835\udc512 ], (2) where Ttig - torque tightening, \ud835\udc53\ud835\udc47 \u2013 friction coefficient on the thread end face; Dm \u2013 mean diameter of ring of contact of nut and bolt ends" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001753_j.ifacol.2019.11.790-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001753_j.ifacol.2019.11.790-Figure2-1.png", + "caption": "Fig. 2. Flat model of the front strut and wheel", + "texts": [ + " (2018a,b); Andrikov et al. (2017). There are various approaches to the description of the tire model, see, for example, Bartram et al. (2010); Sharp and Jones (1981); Pacejka (1980). The vibrations of the wheels in the vertical plane are described well by a simplified two-mass model of the front strut, see Hao et al. (2018). We assume that the longitudinal oscillations of the wheel axis relative to the body are completely absent due to the high rigidity of the corresponding suspension elements (see Fig. 2). The wheel consists of a disk and a tire. The tire is modeled by an external, weightless solid cylinder, which is connected to the disk by numerous weightless springs that are in a pretensioned state. The ends of the springs may be clamped. Similar assumptions about the weightlessness 11th IFAC Symposium on Nonlinear Control Systems Vienna, Austria, Sept. 4-6, 2019 Copyright \u00a9 2019 IFAC 390 A Nonlinear Tire Model to Describe an Unwanted Flat Vibrations of the Wheels Timur V. Glazkov \u2217, Sergey A", + " (2018a,b); Andrikov et al. (2017). There are various approaches to the description of the tire model, see, for example, Bartram et al. (2010); Sharp and Jones (1981); Pacejka (1980). The vibrations of the wheels in the vertical plane are described well by a simplified two-mass model of the front strut, see Hao et al. (2018). We assume that the longitudinal oscillations of the wheel axis relative to the body are completely absent due to the high rigidity of the corresponding suspension elements (see Fig. 2). The wheel consists of a disk and a tire. The tire is modeled by an external, weightless solid cylinder, which is connected to the disk by numerous weightless springs that are in a pretensioned state. The ends of the springs may be clamped. Similar assumptions about the weightlessness 11th IFAC Symposium on Nonlinear Control Systems Vienna, Austria, Sept. 4-6, 2019 Copyright \u00a9 2019 IFAC 390 A Nonlinear Tire Model to Describe an Unwanted Flat Vibrations of the Wheels Timur V. Glazkov \u2217, Sergey A", + " (2018a,b); Andrikov et al. (2017). There are various approaches to the description of the tire model, see, for exam le, Bartram et al. (2010); Sharp and Jones (1981); Pacejka (1980). The vibrations of the wheels in the vertical plane are described well by a simplified two-mass model of the front strut, see Hao et al. (2018). We assume that t e longitudinal oscillations of the wheel axis relative to t e body are completely absent due to the high rigidity of t e corresponding suspension elements (see Fig. 2). The wheel consists of a disk and a tire. The tire is modeled by an external, weightless solid cylinder, which is connected to the disk by numerous weightless springs t at are in a pretensioned state. The ends of the s ri gs may be clamped. Similar assumptions about the weightlessness 11th IFAC Symposium on Nonlinear Control Systems Vienna, Austria, Sept. 4-6, 2019 Copyright \u00a9 2019 IFAC 390 li i l t s i t l t i ti s f t ls i ur . lazkov \u2217, Sergey . esh in \u2217\u2217 \u2217 Ishlinsky Institute for roble s in echanics S, au an osco State echnical niversity, osco , ussia (e- ail: t", + " (2018a,b); Andrikov et al. (2017). There are various ap roaches to the description of the tire model, see, for example, Bartram et al. (2010); Sharp and Jones (1981); Pacejka (1980). The vibrations of the wheels in the vertical plane are described well by a simplified two-mass model of t e front strut, see Hao et al. (2018). We assume that t e longitudinal oscillations of the wheel axis relative to t e body are completely absent due to the high rigidity of the corresponding suspension elements (see Fig. 2). The wheel consists of a disk and a tire. The tire is modeled by an external, weightless solid cylinder, whic is connected to the disk by numerous weightless s ri gs that are in a pretensioned state. The ends of the springs may be clamped. Similar assumptions about the weightlessness 11th IFAC Symposium on Nonlinear Control Systems Vienna, Austria, Sept. 4-6, 2019 Copyright \u00a9 2019 IFAC 390 0x y k m s h b M F FN M Fr Fig. 2. Flat model of the front strut and wheel of wheels are sometimes used in analyzing the motion of vehicles while taking into account the friction, for example, and the stability and instability of the modes under rectilinear motion, see Zhuravlev and Rozenblat (2011a,b). Dissipation inside the tire is not taken into account. Main variables: s is the coordinate of the point of contact between the road and the wheel, x and y are the horizontal and vertical deviations of the center of the tire from the center of the disk, \u03b1 is the angle of rotation of the front wheel disc measured from the vertical, \u03b2 is the angle of the tread of the front tire measured from the vertical, h is the spring length of the front strut", + " In other words, FN = \u2212Fy is the normal force in the case of a touch of the front wheel and the road, FN is equal to zero in the case of detachment of the wheel. The corresponding traction force F has the form: F = fFN sgn ( \u03b2\u0307rb \u2212 s\u0307 \u2212 x\u0307 ) , if \u03b2\u0307rb \u2212 s\u0307 \u2212 x\u0307 = 0; \u2212min ( |Fx|, fFN ) sgn Fx, if \u03b2\u0307rb \u2212 s\u0307 \u2212 x\u0307 = 0, (7) where sgn(\u00b7) is the sign function of a real number, min(\u00b7, \u00b7) is the minimum of two real numbers, f is the coefficient of dry friction, and the expression for Fx is given in (4). Forces F , FN , Fr and torsional moment M\u03d5 (i.e. all forces and moments acting on the tire) and control torque M acting on the disk are shown in Fig. 2. 2.1 Nonlinearity in the Elastic Model of the Tire We assume that the coefficients kr and k\u03d5 (see (1), (3)) are not constant, but depend (as a first approximation) on the generalized coordinates as follows: kr = k0 r + k\u03d5 r (\u03b1 \u2212 \u03b2)2, (8) k\u03d5 = k0 \u03d5 + k\u03d5 r (x2 + y2), (9) where k0 r , k\u03d5 r , and k0 \u03d5 are given positive coefficients. From the relations (8) and (9), it follows that the radial stiffness coefficient kr increases with increasing tire torsion |\u03b1 \u2212 \u03b2| and, conversely, the torsional stiffness coefficient k\u03d5 increases with increasing disk displacement \u221a x2 + y2 with respect to tire", + " In other words, FN = \u2212Fy is the normal force in the case of a touch of the front wheel and the road, FN is equal to zero in the case of detachment of the wheel. The corresponding traction force F has the form: F = fFN sgn ( \u03b2\u0307rb \u2212 s\u0307 \u2212 x\u0307 ) , if \u03b2\u0307rb \u2212 s\u0307 \u2212 x\u0307 = 0; \u2212min ( |Fx|, fFN ) sgn Fx, if \u03b2\u0307rb \u2212 s\u0307 \u2212 x\u0307 = 0, (7) where sgn(\u00b7) is the sign function of a real number, min(\u00b7, \u00b7) is the minimum of two real numbers, f is the coefficient of dry friction, and the expression for Fx is given in (4). Forces F , FN , Fr and torsional moment M\u03d5 (i.e. all forces and moments acting on the tire) and control torque M acting on the disk are shown in Fig. 2. We assume that the coefficients kr and k\u03d5 (see (1), (3)) are not constant, but depend (as a first approximation) on the generalized coordinates as follows: kr = k0 r + k\u03d5 r (\u03b1 \u2212 \u03b2)2, (8) k\u03d5 = k0 \u03d5 + k\u03d5 r (x2 + y2), (9) where k0 r , k\u03d5 r , and k0 \u03d5 are given positive coefficients. From the relations (8) and (9), it follows that the radial stiffness coefficient kr increases with increasing tire torsion |\u03b1 \u2212 \u03b2| and, conversely, the torsional stiffness coefficient k\u03d5 increases with increasing disk displacement \u221a x2 + y2 with respect to tire" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003408_062024-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003408_062024-Figure1-1.png", + "caption": "Figure 1. Structure and coordinate space. Figure 2. Coordinate transformation.", + "texts": [ + " Considering that the excavator working device belongs to the four degrees manipulator and its structure is simple, the inverse kinematics is solved by geometric method in this section. Let the new base coordinate system be ' ' ' ' 0 0 0 0{ , , }X Y ZO , which is derived from the translation of the original base coordinate system 0{ }O to 1{ }O . The coordinates of the bucket tip relative to the new base coordinate system ' 0{ }O can be obtained by geometric relation: 0 1 2 2 2 2 1 2 4 1 2 3 1 2 arctan( / ) arccos[( ) / 2 )] w y x a a a a a (5) In Figure 1, the mutual mapping between the joint space and the actuating space can be acquired according to the geometric relation. 2 2 1 1 1 1 12 2 2 3 2 2 2 2 2 2 24 2 cos 2 cos 2 cos AO BO AO BO AO B CO DO CO DO CO D GF EF GF EF EFG (6) 1 1 1 2 1 2 1 2 3 2 2 3 2 3 3 3 4 3 180 180 AO B BO O AO I O O C O O D CO D O O F FO G GO H O O H (7) ICEMCE 2020 Journal of Physics: Conference Series 1601 (2020) 062024 IOP Publishing doi:10.1088/1742-6596/1601/6/062024 where 2 , 3 and 4 denotes the lengths of the hydraulic actuators for the boom, arm and bucket, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003772_icuas48674.2020.9213924-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003772_icuas48674.2020.9213924-Figure2-1.png", + "caption": "Fig. 2: The mambo parrot micro-MAV schematic, where A is the reference origin, I is the inertial reference frame, and J is the robot\u2019s reference frame.", + "texts": [ + " To minimize the cost function, the predictor sub-block of the NMPC estimates the evolution of the MAV behavior, as well as the behavior of the target. This iterative minimization process is repeated in a cyclic fashion. Finally, the control output in the first step uk,k = u\u2217k = [ \u03c6\u2217k \u03b8\u2217k ]T is sent to the robot. The prediction model is essential to predict the behavior of the system, taking the global given positions as positions of a virtual target. In this section we present the model of a MAV. In this work we present the model of a quadrotor, a MAV with four propellers, as shown in Fig. 2. We will follow the same notation presented by L\u2019Afflitto et al. [3]. In this model, let consider R as a set of real numbers (where Rn and Rn\u00d7m are the n \u00d7 1 real column vectors and the n \u00d7 m matrices, respectively). We assume also that I = [O;X,Y, Z] is the orthogonal, inertial world reference frame, with some origin O. We also consider that J = A;x(t), y(t), z(t), t \u2265 t0, is the robot\u2019s robot frame, which is orthogonal and centered at some point A. Let us assume that the MAV\u2019s center of mass is a point C w", + " The center of mass can vary in time in a generalized model, either when the robot is caring a load or when the center of mass moves due to the displacement of a part of the MAV. Thus, with a, b \u2208 R3, where a = [a1, a2, a3]T , the cross product of a and b is a\u00d7b, in which a\u00d7 \u2206 = 0 \u2212a3 a2 a3 0 \u2212a1 \u2212a2 a1 0 . (1) Note now that I and J, the two reference frames, form two orthogonal bases in R3. A vector a \u2208 R3 is denoted by Ia if it is expressed in I, but if a is expressed in J, then no superscript will be used. As shown in Fig. 2, IFg = mg gives the weight of the MAV of mass m, where g is the gravitational acceleration. Finally, the matrix of inertia of each propeller w.r.t. A is IP \u2208 R3\u00d73, and the matrix of inertia is I \u2208 R3\u00d73. The position of point A w.r.t. the origin O of the inertial reference frame I is rA : [t0,\u221e) \u2192 R3, and the velocity of A w.r.t. I is vA : [t0,\u221e)\u2192 R3. Therefore, the attitude of the MAV frame J with respect to the world frame I is captured by the Euler angles (roll, pitch and yaw), using the rotation 3-2-1 [3]", + " The final cost function (12) is a composition of only two terms. Taking into account the two terms previously described with their weights, the resulting cost function is as follows: J(R\u0302p, U) = Np\u2211 i=1 \u03bb1 \u00d7 |Ir2Dk+i \u2212 I Tpk|+ (12a) Nc\u2211 i=1 \u03bb2 \u00d7 |\u2206Uk+i\u22121|, (12b) where abs(\u00b7) denotes the 1-norm for vector arguments and the absolute value for scalars. Several simulations were performed to validate the modified NMPC controller. However, some comments should be made here: \u2022 In our simulations, we used a Micro multirotor aerial vehicle (Micro-MAV) model (Figure 2) using MatLab/Simulink software. The parameters we use are provided by Andrade [18] in the MatLab model. Their system considers all modules of the real MAV, including but not limited to trajectory generation, sensor fusion, optical flow, a nonlinear model of the MAV, a PD control block (based on the work of Pounds et al. [19]) and a log block. Furthermore, the simulation used the standard Bogacki-Shampine method as numerical solver; \u2022 In our simulations, the desired angular speed is 0.25 rad/s; \u2022 The comparison was performed with the PD position controller based on Pounds et al" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003980_ecce44975.2020.9236100-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003980_ecce44975.2020.9236100-Figure4-1.png", + "caption": "Fig. 4. Distribution of von Misese stress for an axialgap MRC machine with two poles at 100,000 rpm", + "texts": [ + " The C-FRP applied to the rotor has a trade RotorStator Rotating Magnetic Field generated by magnetizing current \u03c91 Coils Capacitor Wireless Power Transfer Rotation \u03c9r Capacitor Rotating Magnetic Field generated by load current s\u03c91 Non-magnetic Material(FRP) Non-magnetic Material(FRP) Resonance at \u03c91 Resonance at s\u03c91 \u03c91 = s\u03c91 + \u03c9r Principle of magnetic resonance coupling (MRC) Analysis model of axial-gap MRC machine Axial-gap MRC machines with different numbers of poles Cylindrical conformations (FRP) Rotor structure (FRP) Cylindrical conformations (FRP) coil 1127 Authorized licensed use limited to: City, University of London. Downloaded on May 18,2021 at 04:56:18 UTC from IEEE Xplore. Restrictions apply. name of T700S, and it was made by Toray. Also, it had a maximum tensile strength in the range of 3500 to 7000 Mpa. Fig. 4 shows the distribution of the von Misese stress at a rotational speed of 100,000 rpm, and it indicates that the maximum stress was 4800 Mpa. It was then confirmed that the rotor made of the C-FRP could withstand a centrifugal force at a rotational speed of 100,000 rpm. V. FREQUENCY AND SLIP FREQUENCY The frequency characteristics are important because the behavior of the MRC machine depends on the frequency. Fig. 5 shows the distribution of the magnetic flux density of the axial-gap MRC machine with the capacitors set for a resonant frequency of 1 kHz" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002365_sii46433.2020.9026258-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002365_sii46433.2020.9026258-Figure3-1.png", + "caption": "Fig. 3. The distance between the base of the manipulator equipped with the gripper and the target", + "texts": [ + ", \u02d9\u03b810) T : joint velocities of the manipulator equipped with the RGB-D camera px \u2032 b , py \u2032 b : the position of the robot base \u03b8b : the orientation of the robot base vx \u2032 b , vy \u2032 b : the velocity of the robot base \u03b8\u0307b : the rotational angular velocity of the robot base \u03b86, \u03b87, ..., \u03b810 are located from the base to the edge of the manipulator likewise \u03b81, \u03b82, ..., \u03b85 as shown in Fig.2 (a). In this section, cost functions corresponding to each task are defined. Then, these functions should be differentiable. 1) Approaching the target As shown in Fig.3, the cost function of approaching the target, Harm(\u03b8) is given by larm(\u03b8) which is the distance between the base of the manipulator equipped with the gripper and the target as follows, Harm(\u03b8) = |larm(\u03b8)\u2212 carm|, (6) where carm is constant such that 0 < carm. Harm(\u03b8) will take becomes smaller as larm(\u03b8) becomes closer to carm. 177 Authorized licensed use limited to: Murdoch University. Downloaded on June 16,2020 at 06:09:26 UTC from IEEE Xplore. Restrictions apply. 2) Collision avoidanceAs shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002296_j.mechmachtheory.2020.103849-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002296_j.mechmachtheory.2020.103849-Figure8-1.png", + "caption": "Fig. 8. Intelligent window a.open b.middle c. closed position.", + "texts": [ + " The tuning parameters in the practical applications presented in this paper can be recalculated to achieve different performance on condition that feasible solution can be computed. 4.2. Intelligent window This window is designed to operate according to the sunlight received by a light sensor. In this case could be employed a rotatory actuator which could be a DC motor. In the last case were employed translational SLE units. In this case, the assembly is composed by angulated SLE. This kind of SLE assemblies can form closed mobile structures [20] , as depicted in Fig. 8 . The dimension of the square window in the closed position is 15 centimeters per side. This assembly has 12 revolute joints, which operate with a dynamic friction coefficient equal to 0.01 when the joints are lubricated. Due to constant use or lack of lubrication, the coefficient increase up to 0.1. Thus we have 0.01 < \u03bc < 0.1. Under the fact that there are no abrupt d changes in weather, the desired performance is to have a settling time t s in the range 0.6 < t s < 2. The overshoot is not a constraint however it is important to avoid high peaks" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002291_23744731.2020.1737489-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002291_23744731.2020.1737489-Figure3-1.png", + "caption": "Fig. 3. Fouling agent injection system.", + "texts": [ + " The entering temperature of the water is controlled to be the same for each test of a given coil, while the mass flow rate is controlled to meet the target heat flux rate for each coil. The air velocity is controlled to the nominal face velocity (provided by the manufacturer, for example) and tests are also done at a rate 25% above and below this rate, to increase the generality of the results. The test apparatus is shown in Figure 1. Fouling agent injection is accomplished using a powder feeder, shown in Figure 2, which uses vibration to uniformly feed the fouling agent over time. The powder feeder is Syntron model 13033. A diffuser, shown in Figure 3, was used at the entrance of the wind tunnel to distribute the injected fouling agent uniformly. The diffuser was added after the initial approach, without a diffuser, resulted in visibly non-uniform distribution. The diffuser is a right frustum with bottom opening dimension of 18.25\u201d 12.25\u201d (46.35 31.12 cm), top opening dimensions of 4.5\u201d 3\u201d (11.43 7.62 cm), and a height of 31\u201d (78.7 cm) between the planes of the openings. A summary of the test method is as follows: 1) Install a new coil in the test apparatus and test it" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001342_012015-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001342_012015-Figure1-1.png", + "caption": "Figure 1. UR5 Robot size drawing Figure 2. Cobot", + "texts": [ + "1088/1757-899X/616/1/012015 by the combination of actual test and simulation experiments. Experiments show that this method can not only optimize the trajectory planning time, shorten the time of robot joint movement, but also can be applied in practice. For friction stir welding test, it can greatly shorten the welding cycle, improve work efficiency and increase the life of the stirring head. UR5 cooperative robot is used in this experiment, which is a 6-DOF robot. According to the specifications, the size diagram (Figure 1) is obtained. According to the D-H parameters (shown in Table 1), the robot model can be constructed, and the forward and inverse solutions and singular values [7] can be obtained.Forward and Inverse kinematics.According to the manual of the Cobot UR5(Figure 2), the DH parameters of the robot in table 1 are obtained. The kinematics principle and Simulation of cooperative robot are analyzed [8]. Combining with MTLAB robot simulation library, the forward and inverse kinematics solutions of cooperative robot can be obtained" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000948_ever.2019.8813582-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000948_ever.2019.8813582-Figure1-1.png", + "caption": "Fig. 1. Cross-section of a 12-slot/10-pole HESFPM machine.", + "texts": [ + " This paper is organized as follows. The phenomenon of the open-circuit DC winding induced voltage in a 12-slot/10-pole HESFPM machine is firstly presented and explained. Then, two methods are comparatively investigated by FE method to reduce the open-circuit DC winding induced voltage, i.e., rotor step skewing and unequal rotor teeth. Finally, the reduction effectiveness of these two methods are comprehensively compared and a conclusion is drawn. A 12-stator-slot and 10-rotor-pole HESFPM machine shown in Fig. 1 [13] is employed for the investigation of the open-circuit DC winding induced voltage in the following sections. This HESFPM machine consists of 12-slot stator, three-phase non-overlapping concentrated AC armature winding, non-overlapping concentrated DC field excitation winding, circumferentially magnetized PMs and 10-pole salient rotor. In addition, an iron flux bridge is incorporated into the machine topology to enhance the flux-control capability. The investigation is based on the globally optimized 12-slot/10-pole HESFPM machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001508_3325291.3325388-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001508_3325291.3325388-Figure5-1.png", + "caption": "Figure 5: Cane with road surface discrimination device", + "texts": [ + " On the other hand, the incident angle at the time of lifting cane is larger to around 15 deg. In Figure 4(c), showing the difference \ud835\udc63\ud835\udc37\ud835\udc5b between the approximated surface and the measurement value of the reflection intensity. In conventional method, the road surface distinction rate was low because the measure used for road surface distinction was only the once. In this experiment, comparing the road surface discrimination rate in the proposed method and the conventional method. The cane was installed road surface distinction device shown in Figure 5. Road surface distinction device was mounted at a height of 55 cm from the road surface. The experimental environment is shown in Figure 6. The subjects were 5 people (A~E) of twenties. The experiment was conducted in a room. Obtain measurement data to PC via serial communications using the Bluetooth of the device. Measurement data is transmitting the value of time, distance, incident angle, reflection intensity, temperature, acceleration. As an experimental method, across the street from the START point in the direction of the measurement the road, get 6 times the cane on the ground" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002315_012062-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002315_012062-Figure1-1.png", + "caption": "Figure 1. Schematic model of roller-induced creep phenomenon.", + "texts": [ + " Several influencing factors have been investigated in recent years [6] [7]. Increasing load, loose fits, thin bearing rings and circumferential load (opposed to point load) amplify the effect while higher bearing clearance, wider bearing rings, more rollers and increased friction in the bearing seat weaken ring creep, to only name the main factors. The mechanism of roller-induced creep is generally described as a \u201ccaterpillar-shaped\u201d deformation of the bearing ring due to radial load. In [8], this is exemplarily shown for a simple plate model. Figure 1 schematically depicts this situation for a bearing outer ring. The upper gear in figure 1 shows the initial position with a radial load acting on the rollers, but without a rotation yet. The bearing outer ring shows the \u201ccaterpillar-shaped\u201d deformation (exaggerated in this figure), leading to a small gap in between rollers. Looking at the lower gear, the rollers have moved by an angle \u03b1 and so have the gaps in between them. Due to the prevailing radial and tangential stress situation in the bearing set, the press-fit is overcome locally by tangential strain which leads to a very small movement \u0394\u03b1 of the outer ring relative to its housing, in this case a gear wheel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001316_etfa.2019.8869248-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001316_etfa.2019.8869248-Figure1-1.png", + "caption": "Fig. 1: Quadrotor schematic", + "texts": [ + " In Section III, the Lagrangian model is extended for a swarm of quadrotors and a conservative interatomic-like potential based control law is proposed in order to achieve auto-formation and collision avoidance. The effectiveness of the proposed control laws is verified by numerical simulations in Section IV. Finally we 978-1-7281-0303-7/19/$31.00 \u00a92019 IEEE 963 conclude the paper in Section V. In this section we introduce a simple quadrotor dynamical model and propose a control law which consists of translational and rotational parts based on Langangian dynamics which are to be extended to swarms in Section III. Fig. 1 illustrates the structure of a single quadrotor, the inertial frame, Euler angles, body frame, and motor directions. Let the absolute position of the quadrotor in Cartesian coordinates be denoted by the vector \u03be = [ x, y, z ]> \u2208 R3 and the vector of Euler angles be denoted by \u03b7 = [ \u03c6, \u03b8, \u03c8 ]> \u2208 R3 where \u03c6, \u03b8, and \u03c8 represent the roll, pitch and yaw angles, which respectively are the rotations around the x, y and z axes of its body frame. We denote by the vector \u03c9 = [ \u03c91, \u03c92, \u03c93, \u03c94 ]> \u2208 R4 the angular velocities of the motors of the quadrotor. They are assumed to rotate in the directions illustrated in Fig. 1. Note that every single motor creates a force fi in a righthanded direction. Furthermore, a torque \u03c4Mi is created from the angular velocity and the acceleration of the motor. Since the latter part is negligible as compared to the former part, we have fi = k\u03c92 i , \u03c4Mi = b\u03c92 i where k is the lift constant and b is the drag constant. Then the overall thrust T can be written as T = \u22114 i=1 fi = k\u03c9>\u03c9, which is along the z-axis of the quadrotor body frame. Let the torques defined in the directions of \u03c6, \u03b8 and \u03c8 be respectively denoted by \u03c4\u03c6, \u03c4\u03b8, and \u03c4\u03c8 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001879_012007-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001879_012007-Figure6-1.png", + "caption": "Figure 6. A force closure grasp, with contact surface normals antiparallel to each other.", + "texts": [ + " Then, we obtain the corresponding surface features of each grasp candidate by using the method described in Section 3.5, and attach a label to each grasp candidate to indicate whether or not the candidate is an effective grasp pose by evaluating a force closure grasp [24] would be obtained if the grippers were to close, that is, each of the surface normals at the contacts between gripper and grasped object is opposite to the finger closing direction v and collinear with the connection line between the contacts (see Figure 6). Unfortunately, it is very tricky to determine whether a grasp candidate is a force closure grasp, because the real-time mesh is noise-containing. For example, the mesh data in the YCB dataset is noise-containing because it is reconstructed from the actual sensor. We \u201csoftening\u201d the force closure condition in order to address the above problems. In particular, we consider the point within 2mm of the contact point between two fingers when the fingers are closed after voxelized the point cloud. When the normal of more than 1/4 point that exists in each contact is less than 20 degree with the finger closing direction v , it can be judged as graspable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002493_042027-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002493_042027-Figure2-1.png", + "caption": "Figure 2 Ellipse model of Hertzian contact", + "texts": [ + " Displacement load is the critical load for the plastic behavior of the model. Hertz made two important assumptions in contact analysis [5]: (1) All deformed bodies are in the range of elastic deformation, and do not exceed the limit of material proportion. (2) The load is perpendicular to the surface, and the effect of surface shear stress is ignored. During the point contact process, Hertz established the following normal stress mathematical model, and the normal stress distribution of the contact ellipse is shown in the Figure 2: Where Q is the normal load, a is the long radius of the stress ellipse, and b is the short radius of the stress ellipse,\u03c3 Is the normal stress of the contact ellipse.It is easy to get the maximum pressure center to appear at the geometric center. During the loading process, due to the friction between the gap and the contact surface, the contact angle between the ball and the raceway changes during the actual loading process. As shown in Figure 3, the maximum stress surface of the ball raceway is misaligned, indicating that the raceway and the ball have shifted from each other during the loading process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003334_speedam48782.2020.9161955-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003334_speedam48782.2020.9161955-Figure5-1.png", + "caption": "Fig. 5: Free body diagram of the test bench setup", + "texts": [ + " In order to implement and test various FOC schemes for the test motor, the load motor is fed by an industrial inverter with an integrated control unit (LUST/LTI ServoOne). The torque measurement shaft (Lorenz Messtechnik D-DR2477) is designed for a nominal torque TTS,N = 10 Nm and a nominal limit speed nTS,lim = 10000 \u00b7 1/min, and has an accuracy of 0.25/100 \u00b7TTS,N = 0.025 Nm. The motors are connected to the torque measurement shaft by metal bellows couplings (R+W BK5-15-67). The relevant free body diagram of the test bench for the torque measurement can be visualized according to Fig. 5. It can be seen that within dynamic operations the measured values of the torque sensor do not equal the set torque values of the test motor. If the torque values of the test motor within TABLE I: Inverter (SEMIKRON Semiteach IGBT), test motor (LUST ASH-22-20K13-000), and load motor (LTi LSN127-1200-40-560) characteristics Symbol Description Values Inverter IN Rated output current 30 A UN Rated output voltage 400 V Test Motor: Induction Motor TN Nominal torque 4.7 Nm IN Rated phase current (star connection) 3", + "5 kW nN Nominal speed 3000 1/min Udc,N DC-link voltage 560 V Load Motor: Permanent Synchronous Motor T0 Standstill torque 12 Nm IN Rated phase current 11.67 A nN Nominal speed 4000 1/min Udc,N DC-link voltage 560 V dynamic operation should be determined, the exact behavior of the load machine, the derivative of the speed regarding time, the rotational inertia of the motors, and the behavior of the couplings have to be known and modeled. Therefore, only steady-state torque measurements are focused. Fig. 5 shows that the measured steady-state torque by the torque sensor Tts has to be modified by the friction torque TTM,d to receive the torque TTM at the test motor: TTM(n) = TTS \u2212 TTM,d(n) . (35) The relevant friction torque signals are recorded while the test motor is switched off and they are plotted in Fig. 6. The torque sensor has a relatively low torsional stiffness (890 Nm / rad [24]) which results in a relatively low natural frequency of the system. In order to determine the natural frequency, a constant speed level is adjusted by the load motor and then the control for both load and test motor is deactivated" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001815_s12206-019-1106-3-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001815_s12206-019-1106-3-Figure4-1.png", + "caption": "Fig. 4. An illustration of an STL part build using multi-direction slicing [79].", + "texts": [ + " As compared to the uni-direction slicing variation capability of building part orientation (multi-axis slicing) provides an interfer- ence-free multi-axis tool-part interaction, a synchronized interaction between the laser tool and the material part each precisely but separately controlled, that boosts the build capability for hollowed or undercut features [69]. The methodology selecting optimal build orientation during process planning has been made either by analyzing 3-D part geometry only [70, 71] or by also analyzing 2-D layers while performing multi-axis slicing depending on the type of AM process needed [72-78]. Fig. 4 illustrates an example of multi-axis slicing for wire based AM process that realizes a multiorientated layer model [79]. Planar (or flat) layer has been long and widely preferred by most AM preprocessors for generating 2-D build data out of the 3-D CAD geometry [80]. In fact, surface inaccuracy of AM part, the notorious and inevitable disadvantage of all AM parts for a long time, is occurred mainly due to the planar layer itself [81]; It does not matter whether those are made from either STL models [82] or free form solid models [83]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001911_ijmmme.2020040104-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001911_ijmmme.2020040104-Figure9-1.png", + "caption": "Figure 9. Bending of soft actuators-curl inward", + "texts": [], + "surrounding_texts": [ + "This\ufeffresearch\ufeffis\ufefffocused\ufeffto\ufeffinvestigate\ufeffand\ufeffevaluate\ufeffthe\ufeffsuitability\ufeffof\ufeffSilastic\ufeffP1\ufeffSilicone\ufeffRTV\ufeffand\ufeff Silicone/MWCNT\ufeffcomposite\ufeffmaterial\ufefffor\ufeffsoft\ufeffactuator\ufeffapplications.\ufeffThe\ufeffdumb\ufeffbell\ufeffsamples\ufeffas\ufeffper\ufeff ASTM\ufeffD412-06a\ufeff(ISO\ufeff37)\ufeffstandard\ufeffare\ufeffcut\ufefffrom\ufeffvulcanized\ufeffsilicone\ufeffRTV\ufeffand\ufeffSilicone/MWCNT\ufeff nanocomposite\ufeffand\ufeffuni\ufeffaxial\ufefftensile\ufefftests\ufeffare\ufeffconducted.\ufeffThe\ufeffOgden\ufeffconstants\ufeffare\ufefffurther\ufeffestimated\ufeff from\ufeffthe\ufeffexperimental\ufeffresults\ufeffand\ufeffused\ufeffin\ufeffFE\ufeffanalyses.\ufeffThe\ufeffanalyses\ufeffare\ufeffconducted\ufeffon\ufeffextension\ufeff actuator,\ufeff contraction\ufeff actuator\ufeff and\ufeff bending\ufeff actuator\ufeff made\ufeff of\ufeff these\ufeff materials.\ufeff The\ufeff strain\ufeff of\ufeff both\ufeff extension\ufeffand\ufeffcontraction\ufeffactuators\ufefflooks\ufeffto\ufeffbe\ufeffnonlinear\ufeffdue\ufeffto\ufeffthe\ufeffmaterial\ufeffnonlinearity.\ufeffIn\ufeffregard\ufeff to\ufeffactuators\ufeffof\ufeffSilastic-P1\ufeffSilicone\ufeffRTV,\ufeffthe\ufeffsudden\ufeffincrease\ufeffin\ufeffthe\ufeffstrain\ufeffis\ufeffobserved,\ufeffas\ufeffits\ufeffstiffness\ufeff is\ufefflower\ufeffthan\ufeffSilicone/MWCNT.\ufeff520\ufeffkPa\ufeffand\ufeff200\ufeffkPa\ufeffare\ufeffthe\ufeffmaximum\ufeffinput\ufeffpressure\ufeffthe\ufeffextension\ufeff actuator\ufeffand\ufeffcontraction\ufeffactuator\ufeffof\ufeffSilicone\ufeffRTV,\ufeffrespectively\ufeffcan\ufeffwithstand.\ufeffIn\ufeffregard\ufeffto\ufeffextension\ufeff actuator\ufeffand\ufeffcontraction\ufeffactuator\ufeffmade\ufeffof\ufeffSilicone/MWCNT,\ufeffboth\ufeffof\ufeff these\ufeffactuators\ufeffare\ufeffable\ufeff to\ufeff withstand\ufeff1000\ufeffkPa.\ufeffThe\ufeffresponse\ufeffof\ufeffthe\ufeffbending\ufeffactuator\ufeffmade\ufeffof\ufeffSilicone\ufeffRTV\ufeffis\ufeff84\u00b0\ufeffand\ufeffSilicone/ MWCNT\ufeffactuator\ufeffis\ufeff128\u00b0.\ufeffIn\ufeffa\ufeffshort,\ufeffit\ufeffis\ufeffconcluded\ufeffthat\ufeffSilicone\ufeffRTV\ufeffis\ufeffbetter\ufeffin\ufeffgetting\ufeffthe\ufeffhigh\ufeff strain\ufeffand\ufefffast\ufeffresponse.\ufeffSilicone/MWCNT\ufeffis\ufeffbetter\ufeffin\ufeffachieving\ufeffhigh\ufeffactuation." + ] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure1-1.png", + "caption": "Figure 1 Three-dimensional model of brake disc", + "texts": [ + " University of Shanghai for Science and Technology Tian Yudong and others combined the finite element modal analysis theory and DASP test results to optimize the brake and reduce a certain weight [8]. 2.1.Geometric model In this paper, the front wheel brake disc of a special vehicle is used as the research object. The brake disc used has a ventilating rib between the two working surfaces as ventilation holes, which can increase the strength and reduce part of the weight. The above model uses a circular plate, The material is HT250 and the yield strength is 250MPa. The following figure (Figure 1) shows the three-dimensional model of the brake disc. 2.2.Meshing Due to the special structure of the disc brake, there are many grooves and chamfers, and there are multiple heat dissipation spokes between the double-layer discs, and the internal structure is more complicated. In order to ensure the quality of meshing and calculation accuracy, some simplifications are made when importing the 3D model. According to the actual size of the model, the selected grid cell size is 5mm. The statics analysis module is established on the ANSYS Workbench software platform, and then the above model is exported from Solidworks to the " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003739_s12239-020-0106-8-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003739_s12239-020-0106-8-Figure1-1.png", + "caption": "Figure 1. Modular deformable tire structure diagram. Figure 2 The main components of modular deformable tire.", + "texts": [ + " And then the dynamic simulations of the tire under the two conditions are carried out in ADAMS. Finally, the relationship between the response curves of the applied load and the target load is compared and the force transmission characteristics of the modular deformable tire is obtained. 2. STRUCTURAL DESIGN OF MODULAR DEFORMABLE TIRE 2.1. Tire Structure The modular deformable tire is mainly composed of five parts: wheel frame, piston, petal-type connecting block, wear-resistant rubber block and metal helical spring, as shown in Figure 1 (Kongshu and Zeng, 2019). The tire body module is composed of petal-shaped connecting blocks and wear-resistant rubber blocks, and each tire body modules is assembled circumferentially to form annular tread. Each spoke is provided with a columntype cavity, and the inner connecting part of the hub and spokes is set as an annular cavity, and the column-type cavity and annular cavity conduct with each other to form an air cavity. The tire body module is stuck into the circular arc groove distributed around the rim, and inserted into the piston in the spokes column-type air cavity, which is the main modular bearing component" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002432_humanoids43949.2019.9035008-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002432_humanoids43949.2019.9035008-Figure3-1.png", + "caption": "Fig. 3. Muscle structure of Kengoro\u2019s legs.", + "texts": [ + " We call it muscle in this paper, and use Dyneema as muscle element in Kengoro. Actuation system by muscle has both advantages and disadvantages due to softness. An advantage is hardware softness to contact the environment, on the other hand, a disadvantage is difficulty in control. Errors in joint angle or end-effector position are cause of difficulties to apply ordinary control methods to tendon-driven humanoid. It is difficult to keep knee-bended posture which often used for the initial posture of humanoid balance control. Muscle structure of Kengoro is shown in Fig.3. As muscle actuator of Kengoro, muscle module was adopted, which we have developed for tendon-driven robots [11]. The muscle module is composed of a motor, motor driver, tension sensor and thermal sensor, and is covered by sheet metals. As the number of muscle modules in legs, 17 is for hip, 10 is for knee, 7 is for ankle, 1 is for toe (foot fingers). Several muscles are multi-counted because there are several bi/multiarticular muscles that actuate several joints at once. Muscle arrangement of Kengoro is decided on the basis of anatomical fidelity to human in the terms of muscle attachment points" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000629_ceit.2018.8751827-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000629_ceit.2018.8751827-Figure2-1.png", + "caption": "Fig. 2. Windings of DSIM", + "texts": [ + " Local defects on the outer, inner raceways and ball cause vibrations whose frequencies f0 , fi and fb , are calculated as follows [5]: Outer raceway: fc=f0= nn 2 fr 1- Db Dc cos \u03b2 (1) Inner raceway: fc=fi= nn 2 fr 1+ Db Dc cos \u03b2 (2) Ball: fc=fb= Dc Db fr 1- Db 2 Dc 2 cos2 \u03b2 (3) Characteristic mechanical vibration frequencies described by (1) to (3) can be seen in the stator current spectrum of the motor [6]. This leads to the following frequencies in stator current: fbf=fs\u00b1k fc (4) I. MATHEMATICAL MODEL OF DSIM The stator of the motor considered in this study has six phases divided into two sets of symmetrical three-phase winding (with phase shift between phases of 120\u00ba) separated by an angle (\u03b1 = 30\u00b0) Fig 2. The rotor is a squirrel cage type; the loops are shown in Fig 3. Multiple coupled circuit modeling for the DSIMs includes six differential voltage equations for the stator windings, Nb+1 differential voltage equations for the rotor meshes, and two mechanical differential equations. Thus, the model can be represented in the matrix form by : Vs1 = Rs1 \u00d7 Is1 + d dt \u03c8s1 (5) Vs2 = Rs2 \u00d7 Is2 + ddt \u03c8s2 (6) Vr = Rr \u00d7 Ir + d dt \u03c8r (7) Where: Vs1 = Vsa1 Vsb1 Vsc1 T (8) Vs2 = Vsa2 Vsb2 Vsc2 T (9) Is1 = Isa1 Isb1 Isc1 T (10) Is2 = Isa2 Isb2 Isc2 T (11) \u03c8s1 = \u03c8sa1 \u03c8sb1 \u03c8sc1 T (12) \u03c8s2 = \u03c8sa2 \u03c8sb2 \u03c8sc2 T (13) Vr = Vr1 Vr2 \u2026" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002301_icisct47635.2019.9011865-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002301_icisct47635.2019.9011865-Figure4-1.png", + "caption": "Fig 4. Use of dependencies in Autodesk Inventor.", + "texts": [ + " The first method is the creation of kinematic pairs by setting the degrees of freedom (Fig 3). To do this, select the desired type of kinematic pair from the menu, select the parts that form the kinematic pair and then connect them together, defining edges, axes, planes, points, etc. of interacting components. After the end of operation, the parts are automatically set to the specified location [5]. The second method is to use assembly dependencies created in the context of the Autodesk Inventor assembly (Figure 4). To do this, select two parts that form a kinematic pair, and activate the existing assembly dependencies For the model to work, it is necessary to set to its nodes the appropriate movements. For this, we can operate with various degrees of freedom of kinematic pairs, setting the necessary displacements, velocities, and accelerations [6]. When all kinematic pairs of the mechanism are created and the movements set, we can view and edit the structure of the mechanism by clicking the Repair Redundancies window (Fig 5)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-FigureA.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-FigureA.2-1.png", + "caption": "Fig. A.2. Manufactured constraint mechanism", + "texts": [ + "1, and it can passively rotate and fit in the hole of the node. When pin 1 fits into the hole of the node, it does not rotate in the direction shown in Fig. A.1 and constrains the contraction of the node. On the other hand, when the pin 1 is in the hole of the node, applying a force more than a certain amount in the extension direction causes the pin 1 to come out of the hole and the constraint is released. In addition, the pin 2 is caught in the L-shaped groove, thereby constraining the linear motion between the nodes. Fig. A.2 shows an overview of the manufactured constraint mechanism. Since the pin housing has complicated shape and it is difficult to manufacture by cutting, we made a prototype using an optical modeling 3D printer. A belt-like rubber is used for the spring element of the pin 1 to generate rotational movement. The pin 1 has a complicated shape and it is difficult to manufacture by cutting too. In addition, since the material used for an optical modeling 3D printer is fragile, the pin 1 was made by metal 3D printing with maraging steel" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000925_b978-0-12-816634-5.00003-0-Figure3.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000925_b978-0-12-816634-5.00003-0-Figure3.10-1.png", + "caption": "Figure 3.10 Binder jet (BJ) built stator part. The material is S4 stainless steel and bronze [31]. Photo courtesy: ExOne, Mike Shepherd.", + "texts": [ + " This clearly illustrates that in AM there are various ways to build a part and a careful consideration of the build strategy and process optimization is essential depending on part geometry, material, and commercial requirements. 64 Science, Technology and Applications of Metals in Additive Manufacturing BJ technologies also allow fine-featured part build up with internal geometries. However, during the subsequent furnace sintering process to enhance the strength, part distortion will occur and must be accounted during design and processing. Fig. 3.10 shows a stator part made of S4 stainless steel and bronze [31]. The part replaced a traditionally manufactured 4140 steel stator that would usually be machined out of a solid 65Comparison of various additive manufacturing technologies blank, resulting in significant cost saving. In addition, S4 stainless steel provides better abrasion resistance and longer life for the component. Some DED technologies, such as EBAM, WAAM, and rapid plasma deposition (RPD), are particularly suited for processing very large components with simple geometries due to their very high deposition rate and the large spot size of the energy source (electron beam, electric arc, plasma, respectively)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002991_educon45650.2020.9125271-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002991_educon45650.2020.9125271-Figure3-1.png", + "caption": "Fig. 3 Understanding Mohr-circles", + "texts": [ + " Unfortunately, the essence of the Mohr-Circle is not well understood by many students, although students are generally able to solve a diverse range of Mohr-Circle problems available in numerous textbooks (see for instance [14]). The reason for not grasping the essence of MohrCircles is because the concept is not explained intuitively in textbooks discussing Mohr-Circles. Focus is more on problem-solving and equations resulting from Mohr-Circles, as opposed to understanding the essence and true significance of Mohr-Circles. An exception to above is the elegant explanation of [12]. Fig. 3 explains the fundamental reason why Mohr-Circles are required. Assuming a two-dimensional body is subjected to the forces P1, P2, P3 and P4. It is of interest to determine the internal stresses within the body at an element as shown with a box in Figure 3(a). An exploratory section must then be passed through the element so that the internal forces on the element can be explored. Initially, a choice of an exploratory section a-a is made as shown in Fig. 3(a). Fig. 3 (b) shows the equilibrium of the cut portion. If the resultant of force P1 and P2 is R, then collinear to R, force E must act on section a-a for equilibrium. Section a-a is deliberately chosen such that E is perpendicular to the cut section, hence only produces a normal stress sigma ( ) on the element. The element is not subjected to any shear stresses ( ) in this instance and therefore by definition a-a must be a principal plane (principal planes are planes with no shear stresses). Fig. 3(c) shows another possible section b-b. In this instance, the element must be subjected to both shear and normal stresses as force E has both normal and shear components (N and T, respectively). This suggests the mind-blowing fact that the stresses at an element located inside a body are not independent but dependent on the choice of the plane the analyst makes (a-a or b-b in this case). An intriguing question is then why section a-a or b-b has been chosen and not any other sections? There is indeed an infinite possibility and any arbitrary section through the element could have been chosen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002791_s11431-020-1569-6-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002791_s11431-020-1569-6-Figure4-1.png", + "caption": "Figure 4 (Color online) Buckling load Pi2 of deployable structure versus the number of elements and deployed degree.", + "texts": [ + " In engineering, these parameters can be appropriately adjusted to optimize the maximum carrying capacity of the deployable structure. Also, when the geometrical properties of bars are determined, the changes in the buckling loads Pi1 and Pi2 with respect to the number of elements and deployed angle are separately shown in Figures 3 and 4. It can be observed from Figures 3 and 4 that the buckling load is influenced by the number of units and deployed angle jointly. In Figure 3, it can be seen that the buckling load Pi1 decreases gradually and approaches zero with the increase of the number of units n. In Figure 4, it shows that some critical loads are positive and some is negative. After performing the analysis, the former can be used to determine the buckling load of the structure, however, the latter is not acceptable and should be ignored. We should look for other units to calculate the buckling load of this structure. Comparing Figure 3 and Figure 4, it can be concluded that the criticality of each unit should be considered in the process of calculating, and the minimum of Pi1 and Pi2 is chosen as the critical load of the based-SLEs deployable structure. Also, the results can be applied to the design of a SLE combination mechanism. After determining the geometric properties of the bars, the appropriate deployment angle and the number of units can make the structure more stable, which is very valuable in engineering. In this section, the numerical results related to deployable structures consisting of different SLEs will be employed to demonstrate the influence of different parameters on the buckling of the structure and investigate the correctness and validity of novel stability model proposed in the previous section, which is verified through the simulation in ANSYS and comparison with the other study", + " It can be seen from Table 1 that the minimum value of the critical load for \u03b3=60\u00b0 and 3 units is Pi1=5.3687\u00d710 5 N, which indicates that the third unit in the deployable structure depicted by Figure 5(a) should be used to determine structural instability load and the AO3 section and CO3 section in the unit are most likely to be unstable. Also, it can be seen that the buckling load corresponding to Pi2 is negative when the number of units is 2 and 3, which are not considered because we made a mathematical substitution P/EI=k2 in Section 2. These results are consistent with those described in Figure 4. In addition, when \u03b3=30\u00b0 and the geometrical properties of bars are determined, the critical load at different units of the deployable structure can also be obtained, the results are shown in Table 2. In Table 2, the minimum value of critical loads for \u03b3=30\u00b0 is Pi2=1.4723\u00d710 6 N, which is the value for unit 1. This shows that GO1 section and HO1 section corresponding to the first unit in the deployable structure depicted in Figure 5(a) are most likely to be unstable and the negative value should be ignored" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003672_s42835-020-00538-y-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003672_s42835-020-00538-y-Figure2-1.png", + "caption": "Fig. 2 Arrangement of stepped PMs", + "texts": [ + " To avoid saturation of the magnetic field, the coil core is made of non-ferromagnetic material. There are three layers of rotor array and each layer has 8 PMs. The angle between adjacent PM poles in the same layer is 45\u00b0, and the PM poles 1 3 in two adjacent layers differ by 30\u00b0. N poles and S poles are alternately arranged in longitude and latitude. The PMs are made of NdFeB which is a kind of rare earth material with high coercive force. The main parameters of the motor are listed in Table\u00a01. The stepped PM of the motor is illustrated in Fig.\u00a02. The rotor shell of the PMSpM studied in this paper is made of aluminium, but the rotor shell has tolerances in the manufacturing process. We use a lathe to machine the rotor radius to 65.1\u00a0mm, and then the rotor radius meets the design size through uniform grinding. Because the NdFeB material is brittle and hard, the top of the PM cannot be polished into a curved surface. A pole cover made of aluminium is designed and placed on the stepped PM. The upper surface of the cover plate is a curved surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002653_012008-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002653_012008-Figure3-1.png", + "caption": "Figure 3. Bicycle geometric model by milliken and milliken", + "texts": [], + "surrounding_texts": [ + "Three-wheeled vehicle is that operates on three wheels. Three-wheeled vehicle has two configurations which are the tadpole configuration and the delta configuration. A tadpole configuration is a threewheeled vehicle which has two wheels upfront and one wheel at the back while the delta configuration has two wheels at the back and one wheel at the front. The dynamics system of a tadpole configuration three-wheeled vehicle could be very complicated. Few matters to be considered are the turning and tilting system, components such as: turning radius, center of gravity, wheel track, etc. This components are then can be measured and determined to ensure the ride safety of the vehicle. For a three-wheeled vehicle to maintain its stability, a tilting system is developed. Tilting was meant to lower the CG so the maximum lateral acceleration that could be taken by the vehicle is higher which results in better stability, and so is the safety. The modelling aims to represent the physics that is happening with the vehicle for future research in active steering and tilting assist for a three-wheeled vehicle." + ] + }, + { + "image_filename": "designv11_80_0000387_978-981-10-4938-5_13-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000387_978-981-10-4938-5_13-Figure3-1.png", + "caption": "Fig. 3 (a) Bottom-up configuration; (b) top-down configuration of a vat photopolymerization machine tool for additive manufacturing", + "texts": [ + " 165 Photopolymer-Based AM Methods 166 Additive manufacturing methods that rely on the selective photo-initiated polymer167 ization of polymer resins held in a resin container are known by numerous propri168 etary names, yet all fall under the denominator vat photopolymerization as given by 169 the ISO/ASTM 52900:2015 naming convention. These processes rely on either the 170 digital-driven scan-line rasterizing of an image by means of a laser point source or on 171 the projection of image masks from a projection source to selectively consolidate 172 the photopolymer to the workpiece. Two primary machine configurations exist: 173 a top-down and a bottom-up configuration as seen in Fig. 3. 174 Metal Powder Consolidation 175 Metal powder consolidation is the underlying working principle for a family of 176 processes by which metal powder is selectively consolidated to the near-net shape of 177 a component by either sintering or melting. The consolidation is predominantly 178 induced by means of a high-powered laser and in an inert atmosphere, yet some 179 metal powder consolidation machine tools operate by means of an electron beam in 180 a high-vacuum atmosphere. It is noticeable that unlike the previously described 181 photopolymerization methods, not only geometry but to a high degree also physical 182 properties of the component can be structured in-process" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003881_aiea51086.2020.00096-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003881_aiea51086.2020.00096-Figure7-1.png", + "caption": "Figure 7. Magnetic particle testing.", + "texts": [ + " The inner and outer chain plates of the main drive chain after fatigue test were tested by penetrant testing and magnetic particle testing respectively. The penetrant test results of main drive chain after fatigue test under no lubrication condition are shown in Fig. 6. It can be seen that there are many rough edges on the inner hole edge of the outer chain plate, which are the initiation ends of fatigue cracks. No cracks were found on the surface of chain plate. The magnetic particle testing results of main drive chain after fatigue test under no lubrication condition are shown in Fig. 7. It can be seen that there is magnetic particle accumulation at the edge of inner hole of inner chain plate, indicating that there is crack near the surface. 430 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on November 13,2020 at 07:04:54 UTC from IEEE Xplore. Restrictions apply. IV. CONCLUSION The fatigue life of the main drive chain under different working conditions is tested and studied in this paper. The following conclusions can be drawn: 1. The breaking load of the new main drive chain and the main drive chain with different service time did not decrease obviously with the increase of service time, and the breaking load met the standard requirements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001303_chicc.2019.8865927-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001303_chicc.2019.8865927-Figure3-1.png", + "caption": "Fig. 3 Spatial geometric relationship of collaborative tracking", + "texts": [], + "surrounding_texts": [ + "The crux of the formation control in leader-follower system lies in to precisely control the distance and yaw angle difference between leader and followers according to the movement state of the leader, which means by working out the guidance law of the followers, the desired formation can be ensured, and the UAV can be kept in a continuous tracking state. In a leader (target) coordinate system, the dynamic equation of UAVs can be expressed as: cos( ) sin( ) sin( ) cos( ) i ix i iy i i ix i iy i L L x v v y v v (3) In the equation, ixv and iyv represent the components of the followers\u2019 velocity along the x and y directions in the body coordinate system of the leader respectively; L and L represent the yaw angle and angular velocity of the leader respectively. Formation control demands that followers and leader maintain the angle constant and distance constant . According to the geometrical relationship: ( ) cos( ) ( )sin( ) ( )sin( ) ( )cos( ) cos sin x L i L L i L y L i L L i L x y x x y y x x y y (4) Define the direction error i Le , take the derivative of x with respect to time, and plug it into the equation of state: cos( ) sin( ) sin( ) cos( ) x y L ix iy Lx y x L iy iy Ly v e v e v v e v e v (5) The formation errors of the whole system are defined as d x x xe , d y y ye , where d x and d y represent the expected distance between followers and leader in the x and y directions in the body coordinate system of the leader. Take the derivative of the formation error and simplify the result: ( ) cos( ) sin( ) ( ) sin( ) cos( ) d x y y L ix iy Lx d y x x L ix iy Ly i L e e v e v e v e e v e v e v e (6) The error dynamic equation of the above three equations can be solved according to the spatial geometry and kinematics equations. For the formation system, it is necessary to design a control law to converge the above error equation to 0 in a finite time. Taking the speed as the control input, the following dynamic error equation can be obtained: = ( ) ( ) X F X G X v (7) where 1 2( ) x ix y L y iy x L i L e e e e e e e e X ,v ,F X (8) 1 2 cos( ) sin( ) 0 ( ) sin( ) cos( ) 0 , 0 0 1 d Lx L y d Ly L x e e v e e v G X (9) The error conversion equation is defined as 2 f dt X + k X , where fk is a constant matrix. If the dynamic error 2 is maintained in the plane 2 0 , then the dynamic error will be 0 regardless of the value of fk . Take the derivative with respect to 2 , and substitute the error dynamic equation, then: 2 ( ) ( ) f 0F X G X v k X = (10) Under ideal circumstances, the dynamic error of the system can meet the requirements of the control objective, and the control input is: 1( )( ( ) )fv G X F X k X (11) In order to verify its effectiveness, dynamic error and more stable robust input can be obtained by expanding its stable region: 2 2 1 2 ( ) ( ) sgn( ) ( )( ( ) sgn( )) f f L v L F X G X v k X = G X F X k X (12) At the same time, the more achievable and effective follower guidance rate can be obtained through the polar coordinate equation of the leader-follower formation model, so that the UAV can maintain the formation structure more accurately and quickly and ensure the tracking effect. The polar coordinate motion model of the followers with the leader (target) as its coordinate origin can be described as follows: cos( ) cos( ) sin( ) sin( ) i i i i L i L i i i L i L i i v v v v (13) Where, represents the position deviation angle of followers relative to the leader, and represents the yaw angle, then the acceleration and angular velocity guidance law of the followers can be expressed as: ( ) 2 ( ) i L i L i i L i L i i L i v i L i L k Nu v k v v v v (14)" + ] + }, + { + "image_filename": "designv11_80_0002624_icrom48714.2019.9071860-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002624_icrom48714.2019.9071860-Figure2-1.png", + "caption": "Fig. 2. a) longitudinal and transversal deflection b) Free body diagram of deformed segment of beam", + "texts": [ + " In order to capture the large deflections, nonlinear model of the EulerBernoulli beam theory is used to derive the equation of motion. Considering in-extensionality condition in the midplane of clamped beams, following relation between longitudinal u and transversal deflection w along the midplane [22]: (1) The equation of motion is derived according to Newton\u2019s method as follows: Deformed and un-deformed beam configurations, longitudinal and transversal displacements along with acting forces is depicted in Fig. 2. Based on Fig. 2, using Newton\u2019s second law of motion and considering trigonometric relations reported in [11], following equations can be achieved: (2) where m and M are respectively mass of a unit length and bending moment which evaluate as: (3) Subscripts MSM and S stands for magnetostrictive and steel materials respectively and \u03c3 and \u03c1 represent stress and mass density. According to the constitutive laws of magnetostrictive and steel, stress is related to the strain as: (4) where ES and EMSM respectively represent Young's modulus of elasticity of the steel and magnetostrictive material in constant magnetic field" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001081_ccdc.2019.8833333-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001081_ccdc.2019.8833333-Figure2-1.png", + "caption": "Fig 2. Structure of two-speed transmission (1 - reducer housing 2 - reducer 3 - transmission housing 4 - 1st order planetary alignment 5 - 2st solar wheel 6 - 2st planet wheel 7 - 2st planet bracket 8 - lining 9 \u2013 2st gear ring 10 - transmission housing 11 - worm gear 12 - wedge slide13 \u2013 worm 14 -motor connection flange)", + "texts": [], + "surrounding_texts": [ + "ENGINEERING VEHICLE POWER SYSTEM XDE110 electric dump truck is a 100-ton mine dump truck driven by ac motor. Based on this model, the electric off-road vehicle with two-speed transmission is designed, and the optimal ratio of the two-speed transmission ratio is determined. The power transmission route of this type of electric off-road vehicle is electric motor, two-speed transmission, main reducer, differential and rear axle. B is the power battery of pure electric vehicles, M is the driving motor, T is the two-speed transmission, F is the main reducer, and D is the differential transmission. The maximum stable speed of the vehicle is 50km/h, while the speed of the dumper is generally lower than 30km/h during operation, which meets the requirements of use. Because this kind of engineering vehicle is mainly used for the short-distance transportation of goods in the mine, and the road conditions in the mine are not very good, so the requirements on the acceleration performance of the vehicle are not high. Table1. Overall design requirements for electric off-road vehicle Performance Indicator requirements Maximum stable speed/km/h 50 minimum turning radius/m 22 maximum slope/% 20 Minimum ground clearance/mm 730" + ] + }, + { + "image_filename": "designv11_80_0000724_978-981-13-3305-7_190-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000724_978-981-13-3305-7_190-Figure3-1.png", + "caption": "Fig. 3. The discontinuous transition strategy", + "texts": [ + " Transition mode is the most important flight stage in the whole flight procedure for tiltwing aircraft, during which the aircraft should reach the required flight velocity rapidly, simultaneously keep a high level of flight stability and performance under the effect of various disturbances and uncertainties, for instance, the large AOA aerodynamic effect, the change of the location of CG, and the outer gust in random direction. Most of these disturbances and uncertainties can be reduced and even eliminated through the method of robust control, except the large AOA aerodynamic effect, which happens in the procedure of the discontinuous transition mode. Fortunately, it can be reduced by changing the discontinuous strategy into continuous transition strategy shown in Fig. 3, which means that the magnitude of AOA is kept roughly the same during the whole transition procedure, meanwhile, the flight path is a continuous curve in the longitudinal plane. However, compared with discontinuous transition strategy, continuous transition strategy shown in Fig. 4 requires higher level of state estimation accuracy and flight performance to maintain the AOA at a constant value. Besides, more effective and powerful propulsion system is needed to obtain the larger tangential acceleration of the flight path to reach the minimum flight speed of fixed-wing mode in a relatively short time" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003709_eit48999.2020.9208299-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003709_eit48999.2020.9208299-Figure3-1.png", + "caption": "Figure 3. PID controller test set up", + "texts": [], + "surrounding_texts": [ + "The newly developed course ECE 410 Cyber Physical System has two components, Physical and Cyber systems presented in Fig. 1, consisting of lectures, and hands-on laboratory assignments. The class was designed in flip class style so that students can use open lab hours to complete hands-on lab assignments. Physical system course topics are summarized in Table 1. A series of formative and summative evaluation tools was used to determine the progress and effectiveness of this proposed CPS curriculum program at periodic intervals equal spaced throughout the CPS project to ultimately determine the impact of the programs on students. External evaluations will be conducted in the second year, 2020. Our evaluation will employ a pre- and post- anonymous survey to gauge the impact of the program, both in terms of participant growth and to determine areas where the program might be improved. Our hope is to utilize existing tools that combine both qualitative and quantitative information on what the students learned and their perspectives how their knowledge base has been broadened in the areas of cyber physical systems. In addition, participants will be polled for ideas and recommendations for program improvements, new lines of inquiry, and suggestions for activities are a vital component of program development. Fig. 2 shows the printed circuit board students will use for the final project of this physical course. The literature survey marks the start of the course instruction and will require the students to venture out and find what state-of-the-art technologies currently exists, and what would be good candidates to accelerate using the techniques they learned. At this point, the students will have a viable topic of interest with source code and resource supplements. The students will have to build baseline hands-on environments to reproduce the results. Through this reproduction, a study will be conducted to analyze and identify likely bottlenecks. A potential way to identify bottlenecks can be a through profiling, which will analyze both memory and CPU, usage and bandwidth, and at different components of the reproduction. Upon identification of a viable bottleneck with appropriate analysis, a design will be enacted to accelerate and improve the current design through the students\u2019 selected method, i.e. CPU, FPGA, or GPU, to define the structure and implementation process. The design will be implemented and tested and if the implementation was successful, the project will be complete, with a paper and poster to follow. The practical session will allow additional time for re-design, implementation, and testing and will be adjustable according to each individual student\u2019s progress. Authorized licensed use limited to: Auckland University of Technology. Downloaded on November 02,2020 at 01:10:29 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_80_0001685_icems.2019.8921850-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001685_icems.2019.8921850-Figure2-1.png", + "caption": "Fig. 2. Finite element and faulty model. (a) The solving model of the electromagnetic field. (b) Faulty model of phase A.", + "texts": [ + " The prototype was modified by the addition of a number of wiring terminals connected to the stator coils and shown in Fig. 1. In order to keep the safety of the system, a resistance was used to limit the short circuit current. The motor parameters are shown in Table I. ITSC faults will change the topology of circuit, and bring a short-circuit loop into the three-phase symmetry windings. According to the motor structure parameters and the faulty situations, a two-dimensional electromagnetic field simulation model was established and shown in Fig. 2. Fig. 2(b) shows the ITSC model of phase A, where S1, S2, \u2026, S16 are the slot This work was supported by the Heilongjiang Province Science Fund for Distinguished Young Scholars (JC2016010). 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE numbers in Fig. 2(a). As the prototype is a 2-pole induction motor, part of the flux will pass through the shaft due to the low rotor current frequency. Therefore, the boundary conditions of shaft are ignored in modeling process. III. TRANSIENT MAGNETIC FIELD CALCULATION The FEM models of the healthy and faulty motor were established respectively, and the variations of magnetic fields pre/post fault have been obtained. The distribution of radial air gap magnetic flux density is shown in Fig. 3. It can be seen from the figure that positive and negative waveforms of the curve are symmetrical", + " Compared with the healthy motor, a higher saturation phenomenon occurs around the faulty slot which is caused by an additional magnetic field generated by the short circuit current. The magnetic flux density distribution of the motor pre/post fault under the rated load is shown in Fig. 6. The overall appearance of the magnetic field is symmetrically distributed in two poles in healthy condition. But in faulty conditions, due to the influence of the additional magnetic field which is generated by the short circuit, localized saturation occurs in the teeth around the faulty slot which is the first slot in Fig. 2(a) and the symmetrical distribution of the magnetic field are destroyed. In addition, this phenomenon deepens with the faulty severity increasing. IV. ELECTROMAGNETIC PERFORMANCES CALCULATION AND ANALYSIS The magnetic field is symmetrically distributed when the stator windings are symmetrical. However, when the ITSC fault occurs, the stator circuit topology will become asymmetry. The specific effects of ITSC fault on the electromagnetic performance will be investigated deeply. The simulated stator current waveform is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001911_ijmmme.2020040104-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001911_ijmmme.2020040104-Figure4-1.png", + "caption": "Figure 4. FE model of braided soft actuators: (a) Contraction type; (b) Extension type; (c) Bending type", + "texts": [], + "surrounding_texts": [ + "Finite\ufeffelement\ufeffanalysis\ufeffis\ufeffan\ufeffapproximation\ufeffnumerical\ufeffmethod\ufeffthrough\ufeffwhich\ufeffthe\ufeffsolution\ufefffor\ufeffthe\ufeff engineering\ufeffproblems\ufeffare\ufefffound.\ufeffAny\ufeffengineering\ufeffdomain\ufeffis\ufeffdiscretized\ufeffinto\ufeffsmall\ufeffsegments\ufeffcalled\ufeff elements.\ufeffEach\ufeffelement\ufeffis\ufeffmade\ufeffof\ufeffa\ufeffset\ufeffof\ufeffnodes\ufeffthrough\ufeffwhich\ufeffthe\ufefffield\ufeffvariable\ufeffis\ufeffmeasured.\ufeffIn\ufeffthe\ufeff design\ufeffof\ufeffactuators,\ufeffthe\ufeffdesign\ufeffengineers\ufeffare\ufeffinterested\ufeffin\ufefftwo\ufeffparameters;\ufeffa)\ufeffdisplacement\ufeff(primary\ufeff variable)\ufeffb)\ufeffstress\ufeffinduced\ufeffon\ufeffthe\ufeffbody\ufeffdue\ufeffto\ufeffinput\ufeffpressure.\ufeffThe\ufeffdisplacement\ufeffindicates\ufeffthe\ufefflimit\ufeffof\ufeff the\ufeffactuation\ufeffin\ufeffa\ufeffparticular\ufeffkind\ufeffof\ufeffthe\ufeffactuator.\ufeffThe\ufefflater\ufeffindicates\ufeffthe\ufeffdesign\ufeffcriterion\ufeffthat\ufeffdecides\ufeff suitability\ufeffof\ufeffthe\ufeffmaterial\ufeffand\ufeffgeometry\ufefffor\ufeffthe\ufeffparticular\ufeffapplication. In\ufeffthe\ufeffcurrent\ufeffanalyses\ufeffof\ufeffcontraction\ufeffand\ufeffextension\ufeffactuators,\ufefflinear\ufeffdisplacement\ufeffis\ufeffthe\ufeffprimary\ufeff variable\ufeffof\ufeffinterest\ufeffand\ufeffstress,\ufeffstrain\ufeffare\ufeffthe\ufeffsecondary\ufeffvariables\ufeffto\ufeffbe\ufeffmeasured.\ufeffIn\ufeffthe\ufeffanalysis\ufeffof\ufeff bending\ufeffactuator,\ufeffthe\ufeffbending\ufeffangle\ufeffis\ufeffthe\ufeffprimary\ufeffvariable\ufeffof\ufeffinterest.\ufeffFigure\ufeff4\ufeffshows\ufeffthe\ufeffbraided\ufeff soft\ufeffactuator\ufeffmodels\ufeffand\ufeffthe\ufeffcorresponding\ufeffangles\ufeffused\ufeffin\ufeffthe\ufeffmodels.\ufeffThe\ufeffother\ufeffparameters\ufeffsuch\ufeff as\ufeffactuator\ufeffthickness,\ufeffchamber\ufeffheight,\ufeffchamber\ufeffdiameter,\ufeffsizes\ufeffof\ufeffactuator\ufefftop\ufeffcap\ufeffand\ufeffbottom\ufeffcap\ufeff as\ufeffwell\ufeffas\ufefffiber\ufeffmaterial\ufeffare\ufeffsimilar\ufefffor\ufeffall\ufefftypes\ufeffof\ufeffthe\ufeffsoft\ufeffactuator\ufeffas\ufefflisted\ufeffin\ufeffTable\ufeff2.\ufeffThe\ufeffbraid\ufeff angles\ufeffof\ufeffvarious\ufeffmodels\ufeffwere\ufeffchosen\ufeffbased\ufeffon\ufeffthe\ufeffbest\ufeffperformance\ufeffshown\ufeffby\ufeffactuators\ufeff(Ili\ufeffet\ufeffal.,\ufeff 2014;\ufeffTakaoka\ufeffet\ufeffal.,\ufeff2013)\ufeffin\ufeffthe\ufeffpast.\ufeffTable\ufeff3\ufeffdepicts\ufeffthe\ufeffbraid\ufeffangles\ufeffused\ufeffin\ufeffvarious\ufeffsoft\ufeffactuator\ufeff models.\ufeffThe\ufeffmaterial\ufeffof\ufeffthe\ufeffrubber\ufefftube\ufeffof\ufeffthe\ufeffactuators\ufeffare\ufeffSilastic-P1\ufeffsilicone\ufeffRTV\ufeffand\ufeffSilicone/ MWCNT\ufeffnanocomposite\ufeffas\ufeffdiscussed\ufeffin\ufeffSection\ufeff2.2.\ufeffThree\ufeffmodels\ufeffcorresponding\ufeffto\ufeffthree\ufeffdifferent\ufeff actuators\ufeffwere\ufeffrendered\ufeffin\ufeffMARC\ufeff(a\ufeffnonlinear\ufeffFE\ufeffanalysis\ufeffsoftware)\ufefffor\ufeffeach\ufeffmaterial.\ufeffThe\ufeffmeshing\ufeff of\ufeffeach\ufeffmodel\ufeffwas\ufeffdone\ufeffwith\ufeff3D\ufeffelements\ufeffand\ufeffTable\ufeff4\ufeffdepicts\ufeffthe\ufeffnumber\ufeffof\ufeffnodes\ufeffand\ufeffelements\ufeff used\ufeffin\ufeffeach\ufeffmodel. Obviously,\ufeffthe\ufeffdeformation\ufeffof\ufeffcontraction\ufeffand\ufeffextension\ufeffactuators\ufeffin\ufeffy-axis\ufeffis\ufeffcomparatively\ufeff larger\ufeffthan\ufeffx-axis\ufeffand\ufeffz-axis.\ufeffFor\ufeffbending\ufefftype\ufeffsoft\ufeffactuator,\ufeffthe\ufeffdisplacement\ufeffin\ufeffx-axis\ufeffand\ufeffy-axis\ufeff (horizontal\ufeffand\ufeffvertical)\ufeffare\ufeffcomparatively\ufeffhigher\ufeffthan\ufeffz-axis.\ufeffThe\ufeffboundary\ufeffconditions\ufeffwere\ufeffset\ufeff accordingly.\ufeffThe\ufeffinput\ufeffpressure\ufeffvarying\ufefffrom\ufeff0\ufeffto\ufeff1000\ufeffkPa\ufeffwas\ufeffapplied\ufeffat\ufeffone\ufeffend\ufeffof\ufeffthe\ufeffactuator\ufeff and\ufeffthe\ufeffrespective\ufeffdeformation\ufeffof\ufeffthe\ufefflinear\ufeffactuators\ufeffand\ufefftrajectory\ufeffof\ufeffthe\ufeffbending\ufeffactuators\ufeffwere\ufeff observed.\ufeffThe\ufeffstrain\ufeffof\ufeff linear\ufeffactuators\ufeffand\ufeffbending\ufeffangle\ufeffof\ufeff the\ufeffbending\ufeffactuators\ufeffwere\ufefffurther\ufeff computed\ufefffrom\ufeffthe\ufeffresults." + ] + }, + { + "image_filename": "designv11_80_0002568_0954407020909242-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002568_0954407020909242-Figure1-1.png", + "caption": "Figure 1. A 16-DOF model and plane structure diagram. DOF: degree of freedom.", + "texts": [], + "surrounding_texts": [ + "The proposed articulated engineering vehicle dynamics model is complex, and the subsystems are coupled and interacted with each other, significantly increasing the non-linearity of the entire system. The following assumptions are made for the model: (1) when the vehicle is running, the unevenness of the road surface remains constant, which is only related to the initial road grade; (2) keep driving at a constant speed during the course of driving; (3) ignore air, wheel roundness changes, and the driver\u2019s impact on the vehicle\u2019s stability; and (4) when the articulation angle is zero, the vehicle is bilaterally symmetrical. In this paper, a 16- DOF articulated engineering vehicle dynamic model considering the longitudinal vibration uxi, lateral vibration uyi, vertical vibration uzi, pitch angle ui, roll angle fi, yaw angle ci, and vertical vibration of four nonsuspended mass ui is established, as shown in Figures 1 and 2. A 16-DOF model of the wheel articulated loader is established using the Newton\u2013Euler method.14 The equations of motion are given below m1 \u20acux1 _uy1 _c1 =Fx1 +Fx2 +Cx \u00f01a\u00de m1 \u20acuy1 + _ux1 _c1 =Fy1 +Fy2 Cy \u00f01b\u00de m1\u20acuz1 = m1g Cz + Fz1 +Fz2\u00f0 \u00de + k1 u1 z1\u00f0 \u00de+ c1 _u1 _z1\u00f0 \u00de+ k1 u2 z2\u00f0 \u00de + c1 _u2 _z2\u00f0 \u00de \u00f01c\u00de Ix1\u20acf1 = Fy1 +Fy2 h1 + Fz1 Fz2\u00f0 \u00de0:5B1 + k1 u1 z1\u00f0 \u00de0:5B1 + c1 _u1 _z1\u00f0 \u00de0:5B1 k2 u2 z2\u00f0 \u00de0:5B1 c2 _u2 _z2\u00f0 \u00de0:5B1 +Cyhc1 My sinu \u00f01d\u00de Iy1\u20acu1 = Fx1 +Fx2\u00f0 \u00deh1 + Fz1 +Fz2\u00f0 \u00delf1 + k1 u1 z1\u00f0 \u00delf1 + c1 _u1 _z1\u00f0 \u00delf1 + k2 u2 z2\u00f0 \u00delf1 + c2 _u2 _z2\u00f0 \u00delf1 +Cxhc1 +Czlf2 My cosu \u00f01e\u00de Iz1\u20acc1 = Fx1 Fx2\u00f0 \u00de0:5B1 + Fy1 +Fy2 lf1 +Cylf2 \u00f01f\u00de The longitudinal, lateral, vertical, roll, pitch, and yaw equations of motion of the rear body are as follows m2 \u20acux2 _uy2 _c2 =Fx3 +Fx4 Cx cosu Cy sinu \u00f02a\u00de m2 \u20acuy2 + _ux2 _c2 =Fy3 +Fy4 Cx sinu+Cy cosu \u00f02b\u00de m2\u20acuz2 = m2g+Cz + Fz3 +Fz4\u00f0 \u00de+ k3 u3 z3\u00f0 \u00de + c3 _u3 _z3\u00f0 \u00de+ k4 u4 z4\u00f0 \u00de + c4 _u4 _z4\u00f0 \u00de \u00f02c\u00de Ix2\u20acf2 = Fy3 +Fy4 h2 + Fz3 Fz4\u00f0 \u00de0:5B2 + k3 u3 z3\u00f0 \u00deB2 + c3 _u3 _z3\u00f0 \u00de0:5B2 k4 u4 z4\u00f0 \u00de0:5B2 c4 _u4 _z4\u00f0 \u00de0:5B2 + Cy cosu Cx sinu hc2 \u00f02d\u00de Iy2\u20acu2 = Fx3 +Fx4\u00f0 \u00deh2 + Fz3 +Fz4\u00f0 \u00delr1 + k3 u3 z3\u00f0 \u00delr1 + c3 _u3 _z3\u00f0 \u00delr1 + k4 u4 z4\u00f0 \u00delr1 + c4 _u4 _z4\u00f0 \u00delr1 Cx cosu+Cy sinu hc2 Czlr2 +My \u00f02e\u00de Iz2\u20acc2 = Fx3 Fx4\u00f0 \u00de0:5B2 + Fy3 +Fy4 lr1 + Cy cosu Cx sinu lr2 \u00f02f\u00de Equations of motion of the four non-suspended masses in the vertical direction (front left, front right, rear left, and rear right, respectively) are as follows mf1\u20acu1 = kt1 q1 u1\u00f0 \u00de+ ct1 _c1 _u1\u00f0 \u00de k1 z1 u1\u00f0 \u00de c1 _z1 _u1\u00f0 \u00de \u00f03a\u00de mf2\u20acu2 = kt2 q2 u2\u00f0 \u00de+ ct2 _c2 _u2\u00f0 \u00de k2 z2 u2\u00f0 \u00de c2 _z2 _u2\u00f0 \u00de \u00f03b\u00de mf3\u20acu3 = kt3 q3 u3\u00f0 \u00de+ ct3 _c3 _u3\u00f0 \u00de k3 z3 u3\u00f0 \u00de c3 _z3 _u3\u00f0 \u00de \u00f03c\u00de mf4\u20acu4 = kt4 q4 u4\u00f0 \u00de+ ct4 _c4 _u4\u00f0 \u00de k4 z4 u4\u00f0 \u00de c4 _z4 _u4\u00f0 \u00de \u00f03d\u00de The generalized coordinates of the 16-DOF model are uT = ux1, uy1, uz1, ux2, uy2, uz2, u1, f,c1, u1,f2,c2, u1, u2, u3, u4 The dynamic equations are rewritten as follows M\u20acX t\u00f0 \u00de+C _X t\u00f0 \u00de+K t\u00f0 \u00deX t\u00f0 \u00de=Fz t\u00f0 \u00de+Fr t\u00f0 \u00de \u00f04\u00de where Fz and Fr represent the engine and road surface excitations, respectively. During the driving process, the external excitations mainly include road and engine excitations that are considered in our dynamic model. In this paper, the pavement power spectral density is used to reflect different levels of pavement roughness, which is given as follows Gq n\u00f0 \u00de=Gq n0\u00f0 \u00de nf n0 W \u00f05a\u00de where nf is the space frequency (m21), n0 is the reference frequency (its value is 0.1m21), Gq(n0) is a pavement roughness factor (m3), and W is the frequency index. Generally, the value of W is 2. Because frequency O = 2pn, equation (5a) can be rewritten as Gq O\u00f0 \u00de=2pw+1Gq n0\u00f0 \u00de O n0 W \u00f05b\u00de When O! 0,Gq(O)! \u2018, the above equation can be written as Gq O\u00f0 \u00de=2pw+1Gq n0\u00f0 \u00de O+Oc n0 W \u00f05c\u00de According to the random vibration theory, the following relationship can be obtained Gq O\u00f0 \u00de= H O\u00f0 \u00dej j2Sw \u00f05d\u00de Through equations (5c) and (5d) and the power spectral density Sw = 1 of white noise W(s), the spatial frequency response function H(O) can be obtained as follows H O\u00f0 \u00de= n0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2pGq n0\u00f0 \u00de p Oc + jO \u00f05e\u00de By equation (5e), the differential equation of road roughness of front wheels can be deduced dqf s\u00f0 \u00de ds +Ocqf s\u00f0 \u00de= n0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2pGq n0\u00f0 \u00de q W s\u00f0 \u00de \u00f05f\u00de Equation (5f) is the stationary process of road roughness in the spatial domain, as follows dqf s\u00f0 \u00de ds = 1 _sc dqf t\u00f0 \u00de dt \u00f05g\u00de When the vehicle is traveling at a non-uniform speed, equation (5g) is substituted into equation (5f), and the following expression can be obtained _qf(t)+ _scOcqf(t)= _scn0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2pGq(n0) p W s(t)\u00bd \u00f05h\u00de However, as W\u00bds(t) is a non-stationary white noise, it cannot be directly used. The covariance of W\u00bds(t) is E W s(t1)\u00bd W s(t2)\u00bd f g= d s t2\u00f0 \u00de s t1\u00f0 \u00de\u00bd = d t2 t1\u00f0 \u00de _s t1\u00f0 \u00de \u00f05i\u00de Then, the stationary white noise W1(t) is used to define the general non-stationary process W1(t)= ffiffi _s p , and the non-stationary process can be expressed as E W1 t1\u00f0 \u00deffiffiffiffiffiffiffiffiffi _s t1\u00f0 \u00de p W1 t2\u00f0 \u00deffiffiffiffiffiffiffiffiffi _s t2\u00f0 \u00de p \" # = d t2 t1\u00f0 \u00de _s t1\u00f0 \u00de \u00f05j\u00de Through equations (5i) and (5j), we can know that the random processes W\u00bds(t) and W1(t)= ffiffiffiffi _sc p have the same covariance, so equation (5h) can be rewritten as equation (6). In our model, the driving speed of the vehicle is not constant, and the front and rear wheels are differently excited by the road surface. The front wheel road excitation is expressed as _qf(t)+ _scOcqf(t)= n0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2pGq(n0)sc p W1(t) \u00f06\u00de There is a hysteresis between the front and rear wheel road excitations due to the distance between wheels. The rear wheel road excitation is expressed as qr(t)= qf(t t0) \u00f07\u00de From equations (6) and (7), the rear wheel road excitation can be expressed as qr(t) = 2 _sc lc _qr(t) _qf(t)+ 2 _sc lc _qf(t) \u00f08\u00de where sc is the longitudinal displacement of the front wheel, Oc represents the cutoff angular frequency of pavement space (Oc =2pnc), and W1(t) denotes the white noise with zero arithmetic mean. Engine excitation including forces and moments generated by the engine is also a major factor affecting the stability of vehicles. In our model, the vertical force Fz and the moment Mx, generated by a four-cylinder inline four-stroke engine that is commonly used in construction vehicles, are given in equations (9a) and (9b) Fz =4mrl v 2pnR 2 cos v pnR t \u00f09a\u00de Mx =4 Mx 1 2mr2 v 2pnR 2 sin v pnR t mr2 v 2pnR 2 l 2 2 sin 2v pnR t + a2 sin v pnR t+u2 + a4 sin 2v pnR t+u4 + \" # \u00f09b\u00de where M is the average gas moment of the piston, m is the equivalent mass of the round-trip motion, r is the rotation radius of crank, n is the variable ratio, R is the wheel radius, l is the ratio of crank length to the length of connecting rod, v is the vehicle speed, ai is the amplitude of overturning moment, and ui is the phase angle of overturning moment. In our model, the front and rear bodies have different driving and rotation speeds. Considering the speed difference of the front and rear bodies, the longitudinal, lateral, and vertical absolute speeds of the rear body centroid can be expressed as follows _ux2 = _ux1 cosu _uy1 lf2 _c1 sinu \u00f010a\u00de _uy2 = _ux1 sinu+ _uy1 lf2 _c1 cosu cos (Df) _uz1 + lf2 _u1 sin (Df) lr2 _c2 \u00f010b\u00de _uz2 = _ux1 sinu+ _uy1 lf2 _c1 cosu sin (Df) + _uz1 + lf2 _u1) cos (Df + lr2 _c2 \u00f010c\u00de where Df denotes the relative roll angle between the front and rear bodies (Df=f1 f2). Because a suspension system is considered in our model, when road surface excitation forces are transmitted to the suspension system through tires, the relative roll angles of the front and rear bodies will change with time. This change will affect the speed of the rear body in three directions (x, y, and z). In order to facilitate research, Df is simplified to be proportional to f1 (Df= kf1). Correspondingly, the above speed formulas for the vehicle rear body can be changed into _ux2 = _ux1 cosu _uy1 lf2 _c1 sinu \u00f011a\u00de _uy2 = _ux1 sinu+ _uy1 lf2 _c1 cosu cos (kf1) _uz1 + lf2 _u1 sin (kf1) lr2 _c2 \u00f011b\u00de _uz2 = _ux1 sinu+ _uy1 lf2 _c1 cosu sin (kf1) + _uz1 + lf2 _u1 cos (kf1)+ lr2 _c2 \u00f011c\u00de Due to the relative rotation of the front and rear bodies, the relationship between the front and rear body yaw angles is given as follows u=c1 c2 \u00f012\u00de Through equations (11a)\u2013(11c), the longitudinal, lateral, and vertical accelerations of the rear body centroid can be expressed as follows \u20acux2 = \u20acux1 _u _uy1 + _ulf2 _c1 cosu \u20acuy1 lf2\u20acc1 + _u _ux1 sinu \u00f013a\u00de \u20acuy2 = \u20acux1 _u( _uy1 + lf2 _c1 sinu+ \u20acuy1 lf2\u20acc1 + _u _ux1 cosu k _f _uz1 + lf2 _u1 cos (kf1) k _f( _ux1 sinu+ _uy1 lf2 _c1 cosu+ \u20acuz1 + lf2\u20acu1 sin (kf1) lr2\u20acc2 \u00f013b\u00de \u20acuy2 = \u20acux1 _u( _uy1 + lf2 _c1 sinu+ \u20acuy1 lf2\u20acc1 + _u _ux1 cosu k _f _uz1 + lf2 _u1 sin (kf1)+ k _f( _ux1 sinu+ _uy1 lf2 _c1 cosu+ \u20acuz1 + lf2\u20acu1 cos (kf1)+ lr2\u20acc2 \u00f013c\u00de In this paper, the vertical displacement of the suspension system is used as the state variable to analyze the force equations, and we are ultimately concerned with the vertical vibrations ux1 and ux2 of the front and rear bodies, respectively, as shown in Figure 3. Therefore, the vertical displacements of four suspensions need state transition. When the front and rear body roll angles f1 and f2 are small, the centroid displacement in the x direction and pitch angle of the front body can be expressed as uz1 = (z1 + z2)=2 and tanf1 =f1 = (z1 z2)=0:5B1, respectively. Similarly, vertical displacements of the four suspensions are obtained as follows z1 = uz1 +0:25B1 tanf1 \u00f014a\u00de z2 = uz2 0:25B1 tanf1 \u00f014b\u00de z3 = uz3 +0:25B2 tanf2 \u00f014c\u00de z4 = uz4 0:25B2 tanf2 \u00f014d\u00de Substituting equations (13) and (14) into the dynamic equation (4) and eliminating Cx, Cy, and Cx related to the hinge, the equations of motion for a 16- DOF articulated loader can be obtained as follows." + ] + }, + { + "image_filename": "designv11_80_0002832_ieeeconf48524.2019.9102530-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002832_ieeeconf48524.2019.9102530-Figure2-1.png", + "caption": "Fig. 2. One step model", + "texts": [ + " By slotting the two steps of slotting in the magnet edge, it could promise to provide a new flux barrier in the magnet surface and optimize the magnet flux distribution in the air gap of the PMMs. II. PMM MODEL PROPOSED The Initial model, one of step model and the two steps model of PMMs studied in this paper as illustrated in Figures 1, 2, and 3, respectively. In Figure 1, it observed that the magnet structure of the PMMs adopted the conventional model. In the model, the height of the magnet was homogenous in all parts of the magnet structure. Thus, the height in the edge of the magnet is the same as the height in the center. Figure 2 shows the one step slot employed in the magnet edge. By using one step slot causes the magnetic flux distribution to become changing and decreasing in the edge of the magnet. Also, the total magnetic flux flowing into the air gap becomes reduced. Figure 3 represents the magnet model of PMM proposed studied in this paper. In this model, a-two steps slot was employed in the magnet edge to achieve and reduce the total magnet flux flowing into the air-gap of the PMM. Using the two steps slot in the magnet edge might expect the reduction of the amplitude of the CT" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000357_pedes.2018.8707615-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000357_pedes.2018.8707615-Figure6-1.png", + "caption": "Fig. 6. Flux Density Dissemination of IM. (a) Healthy Motor. (b) With bearing friction coefficient 0.1 Nm-s/rad. (c) With bearing friction coefficient 0.3 Nm-s/rad.", + "texts": [ + " The distribution of healthy motor is symmetric and the pole magnetic axes are positioned at 90 with respect to each other. The arc of circumference covering all the poles is D/P at any moment of time, where D is the stator internal diameter. When the bearing friction rises, the dissemination becomes unsymmetrical and the flux lines get distorted. The asymmetry escalates with the sternness of fault which is shown in Fig. 5. Hence bearing wear and tear fault creates a deformation in magnetic field and causes a deviation in the magnetic pole axis. Fig. 6 portrays the dissemination of flux density of normal and faulty IM. Compared to the symmetric flux density distribution of normal IM around the poles, the flux density of wear & tear faulty IM is not uniform around the poles which is indicated in Fig. 6(b) & Fig. 6(c). The non uniformity escalates with the sternness of the fault. Fig. 7 indicates the sinusoidal radial airgap flux density of a four pole IM at rated load in both normal & abnormal bearing friction coefficient conditions with respect to the radial distance of the airgap. The radial component of airgap flux density of normal IM contains fundamental component, harmonics of stator MMF & rotor MMF and slot permeance of stator and rotor. Under bearing wear & tear faulty situation the radial component of airgap flux density is severely deformed and unsymmetrical in comparison to healthy machine" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002366_s42405-020-00265-8-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002366_s42405-020-00265-8-Figure2-1.png", + "caption": "Fig. 2 The geometric relations for the vector field guidance [21]", + "texts": [ + " 4, we release the flight test results based on the suggested guidance scheme to verify the performance and the robustness. Before more discussion, we introduce two types of vector fields. Vector field guidance has advantages for trajectory shaping, so that many researchers or engineers utilize it in the realworld. Therefore,we also utilize it to adjust the length of a gliding route to control the arrival altitude at the target window. Lim proposed five types of vector fields in [21], and two of them, VFol and VFch, will be adopted in this paper. These vector fields conjugate the polar coordinates in Fig. 2. r is the radial distance from the origin, and \u03b8 is the clockwise angle from the x-axis. rd is the radius of specific loitering circles, and \u03c7 is the course (or ground track angle) of an aircraft. x- and y-axes are parallel to the northern and the eastern direction of NED coordinates. The first vector field, VFol, is defined as follows: [ r\u0307 r \u03b8\u0307 ] vg\u221a 1 + p2olr 4 [ \u22121 polr2 ] , (1) where pol is the parameter, which determines the configuration of VFol. The sign determines the direction of vortical flow, and the magnitude determines the strength of the flow" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001769_icmect.2019.8932124-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001769_icmect.2019.8932124-Figure15-1.png", + "caption": "Fig. 15. a) Spring-damper load cell model, b) Dynamic force balance for the bottle/plate system.", + "texts": [ + " THE THEORETICAL MODEL A mathematical model was used for the theoretical modelling of the vibratory phenomena of the filling processes described in section III. The model, with lumped parameters, presents one Degree of Freedom (DoF) and time variant mass. In the hypothesis of transversal components of force not significant (as highlighted in session III.B) and of a linear behaviour of load cell and strain gauge, the model considered represents the best compromise between complexity and performance [7]. Referring to Fig. 15.a, the mass m, collected within the container with respect to time t, can be defined as described by (1). (1) M represents the sum of plate, container and load cell masses; \u03bc(t) represents the mass flow rate, which is zero when the tap is closed, constant when it is completely open and with polynomial trend increasing in the opening phase and polynomial decreasing in the closing phase. Differentiating the momentum of the infinitesimal mass m with respect to time, (2) can be written. (2) For the subsystem plate/bottle (Fig. 15.b), considering the elastic and damping contributes introduced by the load cell, (1) and (2), the dynamics is governed by the following second order linear-time-variant differential equation: (3) m(t)g represents the weight force, \u03b7(t)\u03c1Av2 the force impact (\u03b7(t) is a numeric coefficient that takes dissipative effects into account), \u03c1 the fluid density, A the jet section (growing in the tap opening phase), and v the fluid velocity (decreasing in the tap opening phase). Equation (3) can be rewritten in the form of (4), where the natural frequency \u03c9N and the damping ratio \u03be are enlighten: 2 \u2044 (4) with: \u221a (5) The analytical model was implemented in a software application developed in the Delphi environment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000893_iceee2019.2019.00068-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000893_iceee2019.2019.00068-Figure3-1.png", + "caption": "Fig. 3: Quanser AERO configured as a 2-DOF helicopter.", + "texts": [ + " [8] and given by the following relation : (24) where denotes the state vector increments, given by: (25) and is the increment control given by : (26) The constraints on the inputs and the outputs due to the actuator technology, the control system security and the quality desired for the controlled system outputs should be taken into account when solving the optimization problem. In an analytical way, the optimization problem can be written as follows: (27) subject to : (28) where is the vector of the futur increments control, defined as: (29) with, for , the vector is defined as: (30) and, for , the tth increment control is defined as: (31) The Quanser aero system known as 2-DOF helicopter is shown in Fig. 3. The non-linear model of the 2-DOF helicopter model is obtained from Euler-Lagrange method. The input variables are the voltages of main rotor and tail rotor and the output variables are the pitch angle and yaw angle.The plant has two channels and there is an interaction between these channels. In order to reveal in full the plant behavior the system should be considered as multivariable. The physical model of the 2-DOF helicopter is obtained based on the following conventions [13]. 1) The helicopter is horizontal and parallel with the ground when the pitch angle is zero, i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002556_1077546320921048-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002556_1077546320921048-Figure1-1.png", + "caption": "Figure 1. Schematic of a typical linear robotic system.", + "texts": [ + " Consider a linear robot manipulator described by the following differential equation M \u20acx\u00f0t\u00de \u00fe B _x\u00f0t\u00de \u00fe Kx\u00f0t\u00de \u00fe D\u00f0t\u00de \u00bc u\u00f0t\u00de (6) where x\u00f0t\u00de 2R n is the robot positions and u\u00f0t\u00de 2R n is the applied control forces. The matrices M ;B;K 2R n\u00d7n denote the mass, damping, and stiffness matrices of the manipulator, respectively (see Appendix 1). The matricesM , B, and K in equation (6) would be constant for the linear robotic system (Huo and Poo, 2013; Ouyang et al., 2014a). A translational 3-link robot manipulator is typically illustrated in Figure 1. In equation (6), D\u00f0t\u00de 2R n is a combination of the frictions, external disturbances, extra nonlinearities, and model uncertainties. In addition, the term D\u00f0t\u00de denotes the calculation error that arises in the control system. Assumption 1. The system uncertainties and disturbance signals D\u00f0t\u00de are supposed to be bounded as kD\u00f0t\u00dek\u221e \u2264 d. The constant d is also a positive known value. Assumption 2. The reference position xd\u00f0t\u00de and its first and second derivatives (i.e. _xd\u00f0t\u00de and \u20acxd\u00f0t\u00de) are considered to be known and bounded signals" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000815_iemdc.2019.8785385-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000815_iemdc.2019.8785385-Figure1-1.png", + "caption": "Fig. 1. Cross section of 2D-FEM model, visualized with JMAG [2].", + "texts": [ + " This paper introduces a reference single-phase induction machine (SPIM) with a phase shift capacitor in Section II. The SPIM equivalent circuit including all necessary parameters to set up a computer model is developed in Section III followed by a detailed description of the simulation approach in Section IV. Subsequently the parameters chosen for performance comparison, total energy and rotor energy during start-up, are explained. The simulated and measured results are then presented in detail. At the end, conclusions are drawn, and further research directions are suggested. Fig. 1 shows the cross section of the 2D-FEM model. Important parameters are listed in Tables I and II. The winding configuration is shown in Fig. 2. A typical operating cycle involves the acceleration of the anode to a speed that allows sufficient heat distribution on the disc and, after the scan, active braking of the disc. Due to inertia and low friction, nearly constant speed is maintained even when the supply is turned off between these two events. Considering the clinical application, the acceleration time should be minimized to a reasonable degree (<2 s)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002381_sii46433.2020.9026006-FigureA.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002381_sii46433.2020.9026006-FigureA.1-1.png", + "caption": "Fig. A.1. Overview of constraint mechanism", + "texts": [ + " In addition, by integrating the linear mechanism and the bending mechanism, we achieved the extension of two nodes and the contraction of one node while bending, and achieved the linear motion while avoiding the obstacle on the axis of extension. In the future, we plan to research improvement of tip position accuracy to take into consideration the rope sending out length or consider visual feedback control of the tip position using a laser range finder. This paper is based on the results obtained from a project commissioned by the New Energy and Industrial Technology Development Organization (NEDO). APPENDIX Fig. A.1 shows an overview of the designed constraint mechanism. Since this mechanism is attached to the tip of each node of the telescopic structure, the total length of the telescopic structure becomes longer by the thickness of this mechanism and the number of nodes. So, we designed to be nested with the next node and achieved a thin mechanism. The pin 1 always receives a force that rotates in the direction shown in Fig. A.1, and it can passively rotate and fit in the hole of the node. When pin 1 fits into the hole of the node, it does not rotate in the direction shown in Fig. A.1 and constrains the contraction of the node. On the other hand, when the pin 1 is in the hole of the node, applying a force more than a certain amount in the extension direction causes the pin 1 to come out of the hole and the constraint is released. In addition, the pin 2 is caught in the L-shaped groove, thereby constraining the linear motion between the nodes. Fig. A.2 shows an overview of the manufactured constraint mechanism. Since the pin housing has complicated shape and it is difficult to manufacture by cutting, we made a prototype using an optical modeling 3D printer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002904_phm-besancon49106.2020.00042-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002904_phm-besancon49106.2020.00042-Figure1-1.png", + "caption": "Fig. 1. Relationship between preload and angle.", + "texts": [ + " It is impossible to predict how many turns the bolt will rotate during the fitting process. After the connected parts is fitted, the preload begins to show a linear relationship with the rotation angle. During this linear process, the preload F\u2019 CaP 360, where Ca is the rigidity of the bolt and parts system, P is the pitch of the bolt, and is the rotation angle. Continue to rotate the bolt, under the same angle increment, the increment of bolt elongation increases, while the increment of preload decreases, and the relationship between F\u2019 and becomes nonlinear. Fig. 1 shows the relationship between the preload and the angle of rotation during the whole bolt tightening process. Among them, the OA segment is segment is bonding process, the BC segment is linear, yielding begins at point C, and the CD segment is the curve segment after yielding. Torque and angle method can be divided into two steps. Firstly, a threshold torque M0 is set, and the torque is controlled to reach the threshold torque. Reaching the threshold torque indicates that the parts are pressed to each other after several turns of the bolt", + " The functions of the 3 modules in the core of fuzzy controller are: fuzzification module D F converts digital quantity into fuzzy quantity, A* R operates approximate reasoning based on the input fuzzy quantities to obtain a fuzzy quantity U, defuzzification module F D converts fuzzy quantity into digital quantity. A common tightening tool is shown in Fig. 3. It contains a servo motor(including an encoder), a reducer, a torque sensor and a screwdriver. We use the torque sensor to obtain the torque value during the assembly process, and then design the controller to control the motor speed, that is, the rotation angle of the bolt. Then according to the relationship between the pretightening force and the rotation angle shown in Fig. 1, we can Authorized licensed use limited to: University of New South Wales. Downloaded on July 26,2020 at 23:21:41 UTC from IEEE Xplore. Restrictions apply. control the bolt to tighten to the specified pre-tightening force. Firstly, we will use fuzzy controller to control the torque to the threshold torque, and then control the motor to rotate a fixed angle to tighten it to the yield state. This paper focuses on the fuzzy control of threshold torque. The structure of the fuzzy control system is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002189_acit47987.2019.8991028-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002189_acit47987.2019.8991028-Figure7-1.png", + "caption": "FIGURE 7. System principle of EHHT.", + "texts": [ + " Interpolating the displacement of each port into torque formula, then the average torque of each port of electrohydraulic servo plate-inclined plunger hydraulic transformer can be obtained as follows: TA = pA \u00b7 d2zR tan \u03b9 8 \u00b7 sin \u03b1 2 \u00b7 sin \u03b4 (5) TB = pB \u00b7 d2zR tan \u03b9 8 \u00b7 sin \u03b2 2 \u00b7 sin(\u03b4 \u2212 \u03b1 2 \u2212 \u03b2 2 ) (6) TT = pT \u00b7 d2zR tan \u03b9 8 \u00b7 sin \u03b3 2 \u00b7 sin( \u03b1 2 + \u03b3 2 + \u03b4) (7) When the hydraulic transformer is in equilibrium, then the algebraic sum of average torque of the three ports equals zero. After simplifying process, the formula of transformer ratio can be expressed as \u03bb = pB pA = \u2212 sin \u03b12 \u00b7sin \u03b4\u2212 pT pA \u00b7 sin \u03b32 \u00b7sin(\u03b4 + \u03b1 2 + \u03b3 2 ) sin \u03b22 \u00b7 sin(\u03b4 \u2212 \u03b1 2 \u2212 \u03b2 2 ) (8) Plugging numbers into formula 8, a curve can be obtained to solve the transformer ratio of electro-hydraulic servo plateinclined plunger hydraulic transformer, as shown in Fig.7. Fig.7 indicates the transformer ratio increases with the increment of rotation angle when the rotation angle of valve plate is within range from 0\u25e6 to 120\u25e6. The transformer ratio stays zero at 0\u25e6 and changes to one at 60\u25e6. In theory, the transformer ratio becomes infinite at 120\u25e6 and takes the dotted line as the asymptotic line around 120\u25e6. Practically, when the rotation angle of valve plate is 0\u25e6, the displacement at port A is zero, which results in zero transformer ratio; when rotation angle is 60\u25e6, port A and B stay symmetrically versus top and bottom dead point, so that the displacement becomes 8610 VOLUME 4, 2016 equal and transformer ratio achieves 1; when rotation angle is 120\u25e6, the displacement at port B is zero, therefore, the transformer ratio achieves the maximum value. It also can be seen from the curve that transformer ratio and the rotation angle are in non-linear relation, which also accounts for the non-linear characteristic of electro-hydraulic servo piston hydraulic transformer. Fig.7 shows that the transformer ratio changes cyclically with the rotation of the valve plate, with the change cycle of 180\u25e6. That means if Port A and B do translational motion heading right 60\u25e6 along the displacement curve from 0\u25e6 to 120\u25e6, then a displacement curve from 180\u25e6 to 300\u25e6 can be obtained, just with a delayed angle of 180\u25e6. Therefore, if the control angle of the hydraulic transformer changes from 0\u25e6 to 120\u25e6 or from 180\u25e6 to 300\u25e6, a full range of transformer ratio of hydraulic transformer can be realized. III. MODELING OF EHHT The simulation results of EHHT show that when the range of control angle is 0\u25e6 to 100\u25e6, the pressure transformer ratio is 0 to 3, be able tomeet the actual needs so that the swingmotor, which swing angle range is 100\u25e6, was designed to drive the port plat. The schematic of EHHT was shown in Fig.7. The transfer function of servo amplifier is: i = Kau (9) Where i is the output current of amplifier (A), u is the input control voltage of amplifier (V), Ka is the coil circuit gain of amplifier (A/V). Electro-hydraulic servo valve is a crucial control element of electro-hydraulic servo plate-inclined plunger hydraulic transformer. By transforming the minute current signal into hydraulic flow, it can control the swing angle of the hydraulic motor, and thus achieve the control of the hydraulic transformer" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001569_ecce.2019.8912739-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001569_ecce.2019.8912739-Figure6-1.png", + "caption": "Fig. 6 Efforts distribution", + "texts": [ + " G is calculated by dividing the active part into different geometric shapes, then taking the sum of the mass of the shapes, multiplied by their positions, and divided by the total mass of the active part (Fig.5). 2\u03a9\u00d7\u00d7= Glcentrifuga RmF (1) Where RG is the radius of the centroid point G and \u2126 is the rotational speed of the rotor. The FEA of the rotor deformation due to the centrifugal force, shows that: beams 1 and 2 undergo a flexion stress and beam 3 a traction stress. This allows to obtain the effort configuration shown in Fig.6. It has been shown, based on the same previous FEA, that the relationship between the centrifugal force and these efforts is linear and is described by constant coefficients (A, B and C). The calibration between the pre-calculated numerical model and the beam theory model allows the determination of these constants. The analytical equations of the total stress within each beam are summarized in Table 1. The SCF are determined by the calculation of the ratio between numerical stresses calculated on a general geometry of the ASRM, and analytical stresses calculated using the BTM" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003931_icma49215.2020.9233651-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003931_icma49215.2020.9233651-Figure1-1.png", + "caption": "Fig. 1 Structure of the Spherical Amphibious Robot", + "texts": [ + " The feasibility is proved by simulation experiments, and the effectiveness of the algorithm is verified by spherical robot. The rest of this article is organized as follows. In the next section, we will introduce the overall structure of the spherical robot. The third section introduces the method of improved ant colony algorithm, and the fourth section gives the simulation results of the improved algorithm. In the fifth section, the experiment verifies the feasibility of the algorithm. The sixth section gives the conclusion. As shown in Figure 1, the upper half of the robot is made up of a hemispherical waterproof shell made of acrylic material. In order to increase the waterproof performance, the outer part is coated with waterproof glue, and a reasonable space is designed for the placement of sensors. Meanwhile, the structural design is improved to improve the crawling speed and stability of the robot on the land. In order to measure the state of the robot, we install a gyroscope in it, and at the same time. In order to increase the stability of its movement on land, two shaped supports are used to fix the four legs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002405_j.promfg.2019.07.037-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002405_j.promfg.2019.07.037-Figure2-1.png", + "caption": "Fig. 2. (a) design space; (b) contacts; (c) body weight forces; (d) abductor forces; (e) supports; (f) mesh example and (g)(h) mesh details.", + "texts": [ + " Moreover, we also have proved that the final solution with the regular 3D model of the stem has similarities with the real femur, adding one more example to list of best solutions already obtained by nature for a given application. A 3D model with the main femur dimension has been created on Evolve software (SolidThinking) with a simple truncated cone, to have a better understanding of the influence of the bio-inspired 3D model of the stem on the geometrical optimization process. The osteotomy of the major trochanter has been imitated, and only the upper half of the femur has been considered Fig. 2.a. A regular 3D model of a commercial hip stem (Renovis A400 Tappered Cementless) has been created with the software Evolve (SolidThinking). Then, it has been reshaped the regular 3D model on Inspire software (SolidThinking) to include the bone marrow cavity of the femur and this way, getting the bio-inspired 3D model (Fig. 1.b). Other possible bio-inspired shapes to include on the 3D model could have been the Ward triangle (a radiolucent area in the neck of the femur surrounded by principal and secondary stresses)", + " This has been done to check the viability of using a bio-inspired 3D model of the femur based on the assumption that nature already has found a solution close to the optimum one with the femur shape. The bone for the femur model has been simulated as a cortical isotropic bone material and the femoral stem as the titanium alloy Ti6Al4V, suitable for the additive manufacturing of orhtopaedic components [8\u201312]. The material properties for both the bone and the stem are shown in Table 1. The design space, the volume of the stem in which the geometry redesign can be performed, has been all the hip stem, except for two specific areas (Fig. 2.a). These specific areas of the hip stem in which the geometrical redesign has not been performed have been: The tip of the stem, which has been considered as a fully dense material. The Mors-Tapper junction of the stem, thus the junction between the stem and the head is ensured on the final solution. 124 Jaime Orellana et al. / Procedia Manufacturing 41 (2019) 121\u2013128 4 Jaime Orellana/ Procedia Manufacturing 00 (2020) 000\u2013000 Interface constraints between the stem and the bone have been applied to the model. The tip of the femur has been simulated as a sliding contact with no friction, whereas the upper part of the femur has been constrained as a fully bonded contact, a rigid contact, as it can be seen on Fig. 2.b. [13]. Forces due to body weight and the abductor during the gait cycle have been considered, as it can be seen in Fig. 2.c and Fig 2.d. As it is known, during the gait cycle, the forces due to the body weight and due to the abductor change depending on the moment of the cycle considered [14]. Considering all the possible force results for both the body weight and the abductor on the femur would be highly time-consuming for any simulation and studying only one moment of the gait cycle would be not representative. A good compromise between performance and the solution obtained have been found with three load cases obtained during different positions on the gait cycle", + " These three load cases chosen have been considered all together to ensure that the final geometry withstands the complete walking action. The values of these forces can be found in Table 2. Gravity forces due to the weight of the prosthesis have been considered neglectable compared with the forces due to the body weight and the abductor. Similarly, other forces due to body parts (muscles, tendons\u2026) have been neglected for the results here presented as they barely affect the final geometry. Four supports have been applied to the bottom of the femur, as shown in Fig. 2.e. A minimum safety factor of 2 has been imposed for the final solutions. This value is related to the fatigue resistance of the stem and it has been used before in the literature [13, 15] A thickness constraint of 4,0 mm, 4,5 mm, 5 mm, 7,5mm, 10 mm and 15 mm have been applied for both the regular 3D model and the bio-inspired 3D model. This constraint can be understood as a limit for the software to know how precise the geometrical optimization must be. The smaller the minimum thickness (MT) constraint, the more precise it is going to be the geometrical optimization, as the software will be able to remove material up to that MT value", + " 3) The time required to perform all the geometrical optimizations has been obtained from the software timer. Three optimizations have been performed for each model, and then the average time and the standard deviation have been calculated. Errors smaller to 1% on the standard deviation have been neglected. Jaime Orellana et al. / Procedia Manufacturing 41 (2019) 121\u2013128 125 Author name / Procedia Manufacturing 00 (2020) 000\u2013000 5 A tetrahedral mesh has been automatically created by Inspire software. The same mesh generation algorithm has been used for every model. In Fig 2.f-h. it is shown an example of the mesh obtained in one of the models studied. All the modeling, geometrical optimizations, and analysis have been performed on the same computer (MSI GS73 7RE Stealth Pro Intel Core i7), 16Gb RAM, i7, Geoforce GTX 1050 Ti, with a controlled CPU/GPU temperature and with 20% of its maximum memory applied to the simulations, to ensure that all simulations have been done on the same conditions. Applying 100 % of the memory had the risk of interference with any background application" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure5.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure5.5-1.png", + "caption": "Figure 5.5 Scheme of the action of a seismic wave on the lining of a tunnel.", + "texts": [ + " The works [16, 87, 88, 130, 174, 191, 223, 254] are dedicated to the question of interaction of longitudinal or transverse elastic plane waves with the circular openings or inclusions. The calculations of lining of tunnels to seismic resistance is given in the research works [14, 93, 107, 227]. Below we present the methodology of calculations of linings of tunnels on seismic resistance, based on the application of the accelerogram of strong earthquakes. Let us examine a thin-walled ring, which is located in the limitless space; Figure 5.5. It is assumed that the contact between the ring and elastic medium under the seismic exposure (influences) is not disrupted. The seismic action is represented in the form of a plane non-stationary wave. The strength of the ring is possible to check in two limiting (extreme) cases. In the first case the strength of the ring is checked against the effect on a plane longitudinal wave, moving parallel to the axis of ox. In this first case, predominantly vertical oscillations occur on the free surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002630_dese.2019.00015-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002630_dese.2019.00015-Figure4-1.png", + "caption": "Fig. 4. In this subsection ,we hereafter omit the subscript k . The kinematic model of this motion is written as follows:", + "texts": [], + "surrounding_texts": [ + "Lemma 4: Consider the system (10). For any initial condition ( ) [ , ]t\u03c6 \u2032 \u2208 \u2212 , ( ) 0t\u03c6 \u2192 as t \u2192 \u221e . Proof: From section IV-A, ( ) 0t\u03b8 \u2192 as t \u2192 \u221e . If ( ) 0t\u03b8 = , solving (10) gives ( ) exp( ( )) ( )t k t t t\u03c6 \u03c6\u2032 \u2032= \u2212 \u2212 . For arbitrary initial value ( ) [ , ]t\u03c6 \u2032 \u2208 \u2212 , ( )t\u03c6 converges to 0 as t \u2192 \u221e . Moreover, ( ) [ , ]t\u03c6 \u2208 \u2212 always hold as shown in [12]. Therefore, from the converging-input and convergingstate theorem [15], ( ) 0t\u03c6 \u2192 as t \u2192 \u221e . The equilibrium state of bearing angle 0\u03c6 = means that the agent is always aiming at its target. Since the relative positions of follower i and its target j converges, the orientation \u03b7 of follower i converges to the equilibrium state." + ] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure46.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure46.3-1.png", + "caption": "Fig. 46.3 a Calculation of the new rotational pitch angle, ur p;k , due to the run-out on serrated cutters. b Missed cut effect with run-out for serrated cutter", + "texts": [ + " The angular position for the ith flute at height z, measured from the y-axis in a clockwise direction considering the run-out, called the instantaneous radial immersion angle, is calculated as follows ur i z; t\u00f0 \u00de \u00bc Xt\u00fe Xi 1 k\u00bc1 ur p;k 2ztang D \u00f046:4\u00de where X is the clockwise spindle speed (rad/sec) and ur p;k is new pitch angle with respect to Or due to the run-out. The relationship between the new pitch angle ur p;k with respect to Or and the old geometrical pitch angle ug p;k with respect to Og is shown in Fig. 46.3a. From Fig. 46.3a, using the cosine triangle formula, ur p;k is calculated as follows: ur p;k \u00bc cos 1 Rr i \u00f0z\u00de 2 \u00fe Rr i\u00fe 1\u00f0z\u00de 2 Lri \u00f0z\u00de 2 2Rr i \u00f0z\u00deRr i\u00fe 1\u00f0z\u00de \u00f046:5\u00de where Lri \u00f0z\u00de \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Rg i \u00f0z\u00de\u00f0 \u00de2 \u00fe Rg i\u00fe 1\u00f0z\u00de 2 2Rg i \u00f0z\u00deRg i\u00fe 1\u00f0z\u00de cosug p;k q In addition to a change in the instantaneous radial immersion angle and pitch angles, run-out also changes the axial immersion angle, and that is calculated as detailed in [17]", + " Elemental physical static chip thickness is defined as the local distance between previous and current cut surfaces in the direction of normal vector ni\u00f0z\u00de of the flute considering run-out with circular tool path approximation as: hsti z; t\u00f0 \u00de \u00bc gi z; t\u00f0 \u00demin N l\u00bc1 \u00f0Rr i \u00f0z\u00de Rr i\u00fe l\u00f0z\u00de \u00fe fi;l z; t\u00f0 \u00desinur i z; t\u00f0 \u00de sinji\u00f0z\u00de \u00f046:6\u00de where fi;l z; t\u00f0 \u00de is the corresponding feed motion during delay time si;l, ji\u00f0z\u00de is the axial immersion angle, gi z; t\u00f0 \u00de is the screening function due to radial immersion and missed cut (Ref. Fig. 46.3b) effect, explained in [17], and Rr i \u00f0z\u00de is calculated from Eq. (46.3). The missed cut effect considering the run-out creates a non-uniform chip thickness profile as shown schematically in Fig. 46.3(b). Variation in the chip thickness profile for a four-fluted sinusoidal serrated cutter considering zero run-out is shown in Fig. 46.4a, and considering run-out \u00f0q \u00bc 50 lm; d \u00bc 30 \u00de is shown in Fig. 46.4b. The cutting conditions chosen are: depth of cut, ap \u00bc 3mm; feed, ft \u00bc 0:25mm/tooth; and speed, 5000 rpm, for slotting, i.e. 100% immersion. Comparing Fig. 46.4a and b, we see that due to the run-out, the flute chip thickness for the first and second flutes increases by 150%, and for the other two flutes, it decreases by 5%" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000530_978-3-030-12082-5_55-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000530_978-3-030-12082-5_55-Figure9-1.png", + "caption": "Fig. 9 Distribution of heat flux density on contact surface of platform, W/m2", + "texts": [], + "surrounding_texts": [ + "A number of computational experiments were carried out, the conditions and results of which are summarized in Table 2. The temperature field of the macromodel was determined, and then, to compare experiment results, the thermal contact conductance of the gyro unit-platform contact region averaged over the nominal area was calculated by the relation \u03b1c q\u0304 T2 \u2212 T1 where q\u0304 is the average heat flux density over the area of finite elements of the lower nominal surface, T1, T2 are the average temperatures over the area of finite elements of the nominal lower and upper contact surfaces, respectively. Experiments were conducted to evaluate the influence of various factors, while the most appropriate model should be considered as the experiment No. 4 model. The TCC parameter was set as a constant value equal to 157,000 W/(m2 K) or as the above-tabulated dependence on pressure TCC(p). The value 157,000 W/(m2 K) is obtained from the graph in Fig. 5 for the average pressure from the bolt clamp force of 7.3 MPa, calculated by dividing the sum of the clamp forces of each bolt (2000 N) by the nominal contact area. To evaluate the effect of thermal expansion, experiments were carried out for two types of contact behavior, Standard and No separation. For the Standard contact type, the contact heat transfer occurred strictly in the real contact area (Fig. 6), for which the TCC parameter was set. Thus, this type of contact reflects the influence of change in shape from thermal expansion. For the No separation contact type, movement of the contact surfaces along the contact plane is allowed, but separation of the surfaces is not permitted and the real contact area is equal to the nominal one. Thus, in the case of No separation contact, the change in shape of the surfaces from thermal expansion is not reflected in temperature results since it does not affect the thermal contact conductance. In this case, the heat transfer occurs over the entire nominal contact area. The wide use of this type of contact in actual practice is due to the significantly lower computational complexity and, accordingly, solution time. The No separation contact type was set on the gyro unit-platform connection in experiments Nos. 1 and 2. The calculations were carried out with the assumption of small displacements, since the accounting for large displacements for experiment No. 6 resulted in a change in the averaged thermal contact conductance of 0.1%, which is considered insignificant. The first and second experiments set the heat transfer throughout the whole of the nominal contact area. In this case, setting the dependence of the thermal contact conductance TCC on the contact pressure p obtained in the micromodel led to a decrease of 4.3 times in the averaged thermal contact conductance of themacromodel. The models used in the computational experiments Nos. 3 and 4 take into account the effect of thermal expansion on the real contact area. Because of the change in shape of the cylindrical body of the gyro unit, tangency takes place in the form of a narrow ring along the outer edge of the nominal contact area.Also, areas near the bolts are in direct contact. The real contact area was 56% of the nominal area (Fig. 6). As is clear from a comparison of experiments Nos. 3 and 1, the averaged thermal contact conductance decreased bymore than 5 times just due to accounting for the real contact area at a constant TCC of 157,000 W/(m2 K). Under the same conditions and using the TCC(p) dependence (experiments Nos. 2 and 4), the averaged thermal contact conductance decreased noticeably less, by 56%, which can be considered a result of thermal expansion without the direct influence of contact pressure. Repetition of the result of 56% is a random coincidence in this case. Experiment No. 5 showed that the use of a friction coefficient 0.3 instead of 0.5 led to a slight increase in the thermal contact conductance (by 18%). Thus, the friction coefficient has a noticeable effect on the conductance of the actual contact. The thermal contact conductance is significantly affected by clamp force of the bolts. Experiment No. 6 showed that using the conditions of experiment No. 4 and decreasing the clamp force from 2000 to 100 N, the averaged thermal contact conductance decreased by more than 6 times. A similar effect in real structures can arise in the case of more complicated connections, for example, with clasps [35]. The distribution of contact pressures, surface temperatures, and heat fluxes for the contact platform surface in experiment No. 4 is shown in Figs. 7, 8, and 9. Figure 10 shows the distribution of temperatures throughout the entire model of the gyro unit-platform assembly under the same conditions." + ] + }, + { + "image_filename": "designv11_80_0003292_0142331220943071-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003292_0142331220943071-Figure7-1.png", + "caption": "Figure 7. Example 5: (7a) Interconnection of the subsystems at discrete positions. (7b) The circular interconnection of systems.", + "texts": [ + " In the state space form dxc dt (i)= 2m=Ds2 2x a=Ds2 f (i) 2D=Ds2 k(i) xc(i) + m=Ds2 x a=Ds2 0 D=Ds2 xc(i 1) + m=Ds2 x a=Ds2 0 D=Ds2 xc(i+ 1) _xc(i)=Aix(i)+Mi, i 1xc(i 1)+Mi, i+ 1xc(i+ 1), fori= 1, ,N , where xc(i)= a(i, t) c(i, t)\u00bd T . It is assumed that for some i 2 Y the concentration of slime molds can be sensed and for State vectors x(i) of the subsystems can be combined in a large scale system with sparse Ac,Bc and Cc matrices. The subsystems are connected to each other to create a ring shaped structure (Figure 7b). Nonzero structure of the system matrix Ac can be found in Figure 8. The symmetry is broken by scaling M1, 17 by a factor of 0:95 to eliminate poles on the imaginary axis. Continuous time model is discretized by the sampling rate h= 0:01sec. In the simulations, parameters are chosen as Ds= 1 3 10 4,m= 1 3 10 7, x = 4 3 10 4, a= 1:6 3 10 3,D= 3 3 10 8, k(i)= 1:5, f (i)= 0:3, R= I ,Q= I : In the ring, there is an anomalous subsystem that generates cAMP with a higher rate and destabilizes the network" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000020_mwscas.2018.8623843-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000020_mwscas.2018.8623843-Figure5-1.png", + "caption": "Fig. 5. Coordinate systems and forces/moments acting on the quadrotor (m = 1.656 [kg], g = 9.80665 [m/s2], \u2113 = 0.365 [m], I11 = 0.01982 [kg \u00b7m2], I22 = 0.01954 [kg \u00b7m2], I33 = 0.03221 [kg \u00b7m2], kF = 1.79\u00d710\u22127 [N/rpm2], kM = 4.38 \u00d7 10\u22129 [Nm/rpm2])", + "texts": [ + " Flight simulations of a quadrotor vehicle experiencing motor failures In this section, we present typical flight simulations of a quadrotor vehicle experiencing motor failures as shown in Fig. 6\u20139. \u03b8(t) and \u03d5(t) in Fig. 6 and Fig. 8 are completely overlapped. r1(t) and r2(t) in Fig. 7 and Fig. 9 are also completely overlapped. The following results are obtained by Theorem 1\u2013 3 on MATLAB numerical simulations with an ode45 solver. The needed parameters for numerical computations of the quadrotor are shown in Fig. 5. Under the simulation condition in case that the motor in the first rotor is completely failed: \u03b8 = \u03d5 = 0, kF2 = kF3 = kF4 = 1.79 \u00d7 10(\u22127), kM2 = kM3 = kM4 = 4.38 \u00d7 10(\u22129), c = \u22120.24, we have the following flight simulation (\u03c8\u0308 = 12.03512 as the result) achieved by the motor speed control signals of \u03c9M2 = \u03c9M4 = 6652.2505 [rpm], \u03c9M3 = 0 [rpm] as shown in Fig. 6 and Fig. 7. Under the simulation condition in the case that the motors in the first and third rotors are completely failed: \u03b8 = 0, \u03d5 = 0, kF2 = kF4 = 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002766_012022-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002766_012022-Figure1-1.png", + "caption": "Figure 1. The total forces and moments acting on the moving spiral at different D/h ratios with the same volume of WB cells.", + "texts": [ + "1088/1757-899X/826/1/012022 - centrifugal forces \u2211RC directed parallel to the eccentricity line and also perpendicular to the axis of the shaft. They are constant (at \u03c9 = const) in magnitude and also variable in direction; - gas axial forces \u2211RA directed parallel to the shaft axis, variable in magnitude. Gas tangential and centrifugal forces create a tipping moment MOPR acting on the PSP. At the same time, the gas axial forces create a moment opposite to the overturning moment of the sign. We call it the moment of stability My. Obviously, it is necessary (Figure 1): y OPR yK M M , (1) where the stability coefficient KU = 1.1 \u00f7 1.2. In Figure 1(a-b), two possible variants of movable spirals are presented, which provide the same given SPK productivity. Let us compare these options at the same gas pressures, performance, and frequency of orbital motion of the PSP. It can be argued that (with the same scale of drawings): a b OPR OPRM M , (2) a b y yM M . (3) Deflection of the tops of the feathers of spirals: a b PSP PSP . (4) On the other hand: a b A AR R ; (5) RAFT2019 IOP Conf. Series: Materials Science and Engineering 826 (2020) 012022 IOP Publishing doi:10.1088/1757-899X/826/1/012022 a bm m (due to the size of the platform). (6) The dimensions of the thrust bearings, the load on them and the moment of friction in the bearings and the anti-rotation device (PPU) are higher in the second version (Figure 1, b). At the same time, the technological difficulties of ensuring the accuracy of manufacturing spirals (taking into account thermal and force deformation) are more significant in the first version (Figure 1, a). From this list of pros and cons of the two options for spirals, one can see how complicated the issue of optimizing the size of the PSP SPK is. To simplify the problem, we restrict ourselves to considering the stability of the PSP position under the influence of the forces acting on it. We recall that in modern designs of the SPK in the axial direction, the PSP has most often one-sided rigid support - a thrust bearing. The movement of the PSP in the opposite direction is limited only by the force of the gas axial pressure, which does not guarantee it from losing its original position, at least within the gaps, to touching the movable and fixed spirals, which is unacceptable" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003635_052061-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003635_052061-Figure9-1.png", + "caption": "Figure 9. FEM-analyses of segment and contact stiffness", + "texts": [ + " In the present approach the cage is discretized by one node per cage pocket center, displayed in Figure 8 (b). The stiffness matrix /0 of a single segment is simulated by finite elements according to equation 3. /0 12 \u2219 2 3 (3) Vice versa, the contact forces are calculated numerically according to equation 4. Thereby, 34)))\u20d7 is the contact deflection and 5,4 ))))))\u20d7 is the contact force. 5,4 ))))))\u20d7 /0 \u2219 34)))\u20d7 (4) During the contribution only the first entry of the force vector in circumferential direction of the cage is simulated to approximate the stiffness matrix for each SD-element. Figure 9 shows the FEM results of the deformations in circumferential direction 3 that determine the segment stiffness. In case of a rolling element cage pocket contact the stiffness matrix is approximated by the FEM results of the contact pair. The Science of Making Torque from Wind (TORQUE 2020) Journal of Physics: Conference Series 1618 (2020) 052061 IOP Publishing doi:10.1088/1742-6596/1618/5/052061 The global FEM model of the bearing cage aims to show the cage stress as a result of contact forces. The cage model consists of 139 cage pockets that are connected along the axial guiding clearance to the cylindrical system of inertia with one tangential DOF and three rotational DOF" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure40.7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure40.7-1.png", + "caption": "Fig. 40.7 Hot tear region of Al\u2013Cu cast alloy: 2D-top view and 3D-model", + "texts": [], + "surrounding_texts": [ + "Casting simulation was implemented for Al\u2013Cu alloy with green sand mold which enables to predict the problems in casting process in foundry industries. With a motive to minimize the defects in casting especially for green sand molds to acquire the quality of the castings 1. The present study enabled to identify the hot spot and hot tears in casting part by casting simulation. 2. It facilitated optimization of design and casting yield for the preceding process to overcome the hot spot and hot tears by providing feed aids and redesign of the mold cavity. 3. It was concluded that casting simulation enables for trial-and-error to get the accuracy in casting by optimizing the defects in casting, which saves the time and loss of material in foundry which ultimately impacts to enhance the productivity of the industry. 4. It was observed that the Al\u2013Cu alloy in the green sand mold system, the hot tearing and hot spot occurred in the range of 600\u2013700 \u00b0C as the solidification time increases at 120\u2013140 s with a minor variation in the hot tearing and hot spot hence, it was concluded that the hot spot increases, respectively, as hot tearing increases vice versa. Acknowledgements Authors are much thankful for the TEQUIP-III, NPIU-New Delhi, for financial support for purchase of AutoCAST software to carry out the simulation of designed casting. 466 K. S. Pulisheru and A. K. Birru" + ] + }, + { + "image_filename": "designv11_80_0003332_itec48692.2020.9161707-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003332_itec48692.2020.9161707-Figure3-1.png", + "caption": "Fig. 3. Flux density distribution at 3000-rpm, (a) No-load operation, (b) Fullload operation", + "texts": [ + " The DSDR-IPMSynM is modelled for different load currents with various current vector angles to evaluate the steady-state electromagnetic torque. Figs.3 (a) and (b) show the flux density distribution plots under no-load and full-load conditions, for magnetic-static solutions. On the other hand Figs 4 (a) and (b) show the airgap profiles and their harmonic distributions respectively. 463 Authorized licensed use limited to: University of Prince Edward Island. Downloaded on September 08,2020 at 14:02:20 UTC from IEEE Xplore. Restrictions apply. It is noted from Fig.3 (b) that the flux linked with load current increases the flux density in stator and rotor cores of both outer and inner motors, producing high-localized magnetic saturation especially on the stator back iron of the outer motor. On no-load, the d-axis flux density due to pieces of NdFeB35 permanent magnet saturates the tangential bridges of the rotor of the inner motor, which is not the case when the DSDRIPMSynM operates on full-load. Furthermore, the results in Fig. 4 (b) evidence that the outer motor mitigates most of the field harmonics, except the 5th and the 23rd harmonic orders", + " 6 (b) gives the comparison of electromagnetic torque profiles as function of rotor angular displacement, while Table II depicts the values of average torque, torque ripple and its harmonic components. From Fig. 6 (a), it is well noted that under full-load operation, the inner motor achieved a peak torque of about 109 Nm at the current vector angle of 45o elec. On the other hand, the peak torque achieved on full-load by the outer motor is about 299 Nm at the current vector angle of 50o elec. There is an angular shift difference of 5o elect between the inner motor and outer motor achieved peak torques. The difference is due to difference in magnetic saturation levels as noted in Fig. 3 (b). The localized saturation on the outer stator back iron is responsible for shifting the current vector angle by a margin of 5o elect. Noting the results in Table II, it is evident that the inner motor has high torque density per volume; this is due to the merit of concentrated modular winding employed for the inner stator. The outer motor that provides the principal traction to the rear wheels produces 299.1 Nm at the speed of 3000 rpm for a full-load current of 80 A and current vector angle of 50o elec" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003446_jae-200020-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003446_jae-200020-Figure7-1.png", + "caption": "Fig. 7. Model of PM heat conduction, emphasizing slot effect.", + "texts": [ + " (13) un cor rec ted pro of ver sio n Although the slot effect exists, the stator teeth can be still seen as cylinder, but a correction coefficient should be considered. The thermal resistance of stator teeth is calculated as \ud835\udc45\ud835\udc61\u210e_\ud835\udc46\ud835\udc47 _1 = \ud835\udc59\ud835\udc5b(\ud835\udc45\ud835\udc5c\ud835\udc62\ud835\udc61/\ud835\udc45\ud835\udc4e\ud835\udc63\ud835\udc54) 2\ud835\udf0b\ud835\udc58\ud835\udc3f\ud835\udf06\ud835\udc5d\ud835\udc56\ud835\udc5f (14) \ud835\udc45\ud835\udc61\u210e_\ud835\udc46\ud835\udc47 _2 = \ud835\udc59\ud835\udc5b(\ud835\udc45\ud835\udc4e\ud835\udc63\ud835\udc54/\ud835\udc45\ud835\udc56\ud835\udc5b) 2\ud835\udf0b\ud835\udc58\ud835\udc3f\ud835\udf06\ud835\udc5d\ud835\udc56\ud835\udc5f (15) where pir is the correction factor, which is defined as the volume of the stator teeth compare to that of the stator teeth plus slot at a pole distance. Likewise, due to the similar structure, calculation process of permanent magnet resistance is similar. Figure\u00a07 shows the equivalent calculation process of permanent magnet. Due to manufacturing constrains, the air slots exist between permanent magnets. Consequently the coefficient pPM should be considered, which is defined as the volume of the PMs compare to that of the PMs plus slot. The thermal resistance of permanent magnet can be calculated as \ud835\udc45\ud835\udc61\u210e_\ud835\udc43 \ud835\udc40_1 = \ud835\udc59\ud835\udc5b(\ud835\udc45\ud835\udc5c\ud835\udc62\ud835\udc61/\ud835\udc45\ud835\udc4e\ud835\udc63\ud835\udc54) 2\ud835\udf0b\ud835\udc58\ud835\udc3f\ud835\udf06\ud835\udc5d\ud835\udc43 \ud835\udc40 (16) \ud835\udc45\ud835\udc61\u210e_\ud835\udc43 \ud835\udc40_2 = \ud835\udc59\ud835\udc5b(\ud835\udc45\ud835\udc4e\ud835\udc63\ud835\udc54/\ud835\udc45\ud835\udc56\ud835\udc5b) 2\ud835\udf0b\ud835\udc58\ud835\udc3f\ud835\udf06\ud835\udc5d\ud835\udc43 \ud835\udc40 . (17) As to resistance calculation between stator teeth and can, due to slot effect that impedes the heat flow between two adjacent objects, the coefficient CCR should be considered to amend the deviation that caused by the slot effect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure3.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure3.2-1.png", + "caption": "Fig. 3.2 Dimensions for plane\u2013strain tension test", + "texts": [], + "surrounding_texts": [ + "The material selected for this study was high strength low alloy or micro-alloyed steel, EN 10149-2 steel sheet which has 2.42 mm thickness. Chemical composition of material is shown in Table 3.1.\nMechanical Properties and Micro-Structure Evaluation of EN-10149-2\nThe microstructure of EN-10149-2 steel sheet was studied by optical microscope of model Huvitz HR-300 series. Hardness was measured using Vickers\u2019s hardness testing machine. The mechanical properties of EN-10149-2 steel sheet were measured using tensile tests as per ASTM- B557 M by cutting sheet in different rolling directions such as 0\u00b0, 45\u00b0, and 90\u00b0 and evaluated the mechanical properties and plastic strain ratio (R) as per ASTM- E517 through tensile tests. Tensile tests were performed at a nominal cross-head speed of 1 mm/min, at room temperature. On an ISO 68982-1 universal testing machine and repeated thrice for each set to check the reproducibility. Failure from grip or slippage during testing was not observed. Figure 3.1 shows the base metal tensile specimen. The mechanical properties like yield strength, ultimate tensile strength, uniform elongation, strain hardening coefficient (n) and strength coefficient (K) were evaluated as per the standard procedure after testing the samples till failure. The load-stroke behavior obtained during testing and converted into engineering stress\u2013strain and true stress\u2013strain\nTable 3.1 EN-10149-2 (S700mc) steel chemical composition (in wt%)\nC Si Mn P S Cr Mo Ni Cu W Al Nb Ti Fe\n0.06 0.04 1.46 0.003 0.002 0.01 0.02 0.03 0.01 0.04 0.05 0.03 0.10 98.15\nFig. 3.1 Specimens prepared for tensile tests\n3 Formability Evaluation of EN-10149-2 (S700mc) Steel Under \u2026 27", + "plot for evaluating the mechanical properties. The plastic strain ratios at different rolling directions were obtained after testing the base material to 10% plastic strain.\nPlane\u2013Strain Tension Testing\nIn this study, plane stretching testing samples were prepared based on the width constraints method. Figures 3.2 and 3.3 show the schematic and original in-plane plane-stretching test specimens with circular grids of 3 mm in diameters on experiment specimen.\nFigures 3.4 and 3.5 are shown the material EN-10149-2 steel uniaxial tensile test specimen before and after the testing, respectively. Figure 3.5 representing the tested specimen at different rolling direction 0\u00b0, 45\u00b0, and 90\u00b0.\n28 D. H/Georgis et al.", + "The mechanical properties of the EN-10149-2 sheet metal are evaluated from the uniaxial tensile tests tabulated in Table 3.2. The evaluated mechanical properties like yield stress, uniform elongation, tensile stress, material strength coefficient (K) and strain hardening coefficient (n). The base material engineering stress\u2013strain behavior in different rolling directions is shown in Fig. 3.6. Table 3.2 shows the tensile properties of all three reputations and Table 3.3 summarizes the mechanical properties of EN-10149-2 base material. The yield strength and ultimate tensile strength are found to be within 972\u20131020 MPa and 1018\u20131054 MPa, respectively. The uniform elongation and strain hardening exponent are about 6.8% and 0.22 respectively.\nTable 3.4 indicates summarized mechanical properties of EN-10149-2 (S700mc) sheet steel. The mechanical properties and anisotropy value are shown in Table 3.4.\n3 Formability Evaluation of EN-10149-2 (S700mc) Steel Under \u2026 29" + ] + }, + { + "image_filename": "designv11_80_0001462_amm.895.52-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001462_amm.895.52-Figure1-1.png", + "caption": "Fig. 1 Schematic of the Tool-chip tribometer [9]", + "texts": [ + " A regression neural network model with Keras library for Python will be introduced along with the technical aspects of the implementation. Unique equipment called Tool-chip tribometer is used, where experiments were conducted under a pool of oil samples which facilitates the formation of a surface layer similar to the one generated in metal cutting operations. As it gets generated, the layer so formed is friction tested, before it gets exposed. The schematic representation of tool-chip tribometer is shown in figure 1. The friction tests using tool-chip tribometer on AA 6061 are conducted for mineral oil (Servocut 945), raw and modified vegetable oil modes of operations. The test conditions are tabulated in Table 1. The variation in the frictional coefficient during sliding, cutting and sliding after cutting for different speed (100 rpm, 150 rpm, and 200 rpm) at a normal load of 10N was observed and recorded. Artificial neural networks are able to learn and generalize from experience and have powerful pattern recognition and regression capabilities" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001976_s00202-020-00924-9-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001976_s00202-020-00924-9-Figure1-1.png", + "caption": "Fig. 1 Structure of the DSIM windings", + "texts": [ + "com Khoudir Marouani marouani_khoudir@yahoo.fr Azeddine Houari Azeddine.Houari@univ-nantes.fr Mohamed Fouad Benkhoris mohamed-fouad.benkhoris@univ-nantes.fr 1 Ecole Militaire Polytechnique, LSEE UER-ELT, 16046\u00a0Bordj\u00a0El-Bahri,\u00a0Algiers, Algeria 2 IREENA Laboratory, Universit\u00e9 de Nantes, Nantes, France 1 3 The drive system studied in this work is based on a DSIM, where two symmetrical and identical sets of three-phase windings share a common stator magnetic core and are phase shifted spatially by 30\u00b0 (electrical degrees) as depicted in Fig.\u00a01. A voltage source inverter (VSI) is used to feed each set of the stator winding. The rotor of the DSIM is identical to that of the three-phase squirrel cage induction machine. Furthermore, the DSIM is considered as the combination of two three-phase machines sharing the same magnetic core. Hence, the usual Park transformation can be applied to each stator set [8, 13, 14]. Using the vector space decomposition approach introduced in [15], the original six-phase dimensional system can be decomposed into two main sub-models named ( sd1, sq1 ) and ( sd2, sq2 ) for the stator side and (rd, rq) for the rotor side, respectively [16]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003672_s42835-020-00538-y-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003672_s42835-020-00538-y-Figure1-1.png", + "caption": "Fig. 1 Basic structure of the PMSpM", + "texts": [ + " Secondly, the electromagnetic force of the stator coil is calculated based on the Lorentz force method, so that the analytical model of the electromagnetic torque of the motor is constructed. Finally, the AM is compared with the FEM and experimental tests. The rest of the paper is organized as follows. Section\u00a02 introduces the structure of the PMSpM. Section\u00a03 formulates the 3-D magnetic field modelling and FEM verification of the motor. Section\u00a04 illustrates the torque modelling and FEM verification of the PMSpM. Section\u00a05 presents the experiments. Finally, the paper is concluded in Sect.\u00a06. The basic structure of the PMSpM is shown in Fig.\u00a01. It is composed of a stator and a rotor. The former has 24 stator coils and the latter has 24 stepped PM poles. The maximum tilting angle is \u00b1 22.5\u00b0. All stator coils are mounted on the stator shell in two layers and the angle between the coil and the equator is 22.5\u00b0. To avoid saturation of the magnetic field, the coil core is made of non-ferromagnetic material. There are three layers of rotor array and each layer has 8 PMs. The angle between adjacent PM poles in the same layer is 45\u00b0, and the PM poles 1 3 in two adjacent layers differ by 30\u00b0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002253_1350650120908116-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002253_1350650120908116-Figure1-1.png", + "caption": "Figure 1. Bearing and gear shaft structure.", + "texts": [ + " Moreover, there are fewer literature on the investigation of serious surface asperities structure effect on bearings startup characteristics. In this paper, we report the development of a model for the rough journal bearings in external gear pump taking into account the start-up rotation speed variation. An extensive set of simulations have been presented to investigate the effects of the pump operation condition and surface asperities structure on the pump bearings performance during the start-up. Figure 1 shows a schematic diagram of gear pump journal bearings with the main load acting on the shaft. The shaft is rotating with a speed and the bush is fixed. O and O0 represent the shaft and bush center, respectively. The radius direction gap is C. The bush radius and width are R and L, respectively. Oli film thickness in bearings clearance. To obtain more accurate lubrication characteristics in the mixed-lubrication regime, the film deformation caused by surface roughness should not be neglected. The fluid region of bearing with a high eccentricity is presented in Figure 2 with the coordinate system" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002950_042044-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002950_042044-Figure1-1.png", + "caption": "Fig 1. Structure diagram of tubular permanent magnet coupler.", + "texts": [ + " The analysis results lay a foundation for the subsequent analysis of the thermomagnetic coupling of the permanent magnet couplings. Based on the magnetic field, the permanent magnet coupling uses the non-contact relative motion between the permanent magnet and the conductor to form the induced magnetic field. The interaction between the induced magnetic field and the original magnetic field produces the torque and realizes the torque transmit between the motor and the load section. The structure of the tubular permanent magnet coupler is shown in the Fig. 1, which is mainly made of inner and outer steel frames, permanent magnet blocks, aluminium yokes and conductor tubes. The magnetism of permanent magnet is directly related to the temperature, and the meaning high temperature will cause demagnetization, and at the same time the material performance parameters of components of permanent magnet coupler are also related to the temperature, such as the conductivity of conductor barrel, the remanence of permanent magnet, the coercive force of permanent magnet and the permeability of outer steel disk" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002296_j.mechmachtheory.2020.103849-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002296_j.mechmachtheory.2020.103849-Figure10-1.png", + "caption": "Fig. 10. Folding roof: a. folding state b. isometric view c. front view d. top view of deploying state.", + "texts": [ + " Folding roof This model of mechanism can be employed in cars, motorcycles, bikes as well in residential applications for example a balcony, a pool, or a playground. This structure employs spherical SLE [4] . The dimensions of the folding roof in deploying conditions are 30 centimeters of length, 12 centimeters of width and height equal to 15 centimeters. This structure has 14 joints operating with a friction coefficient in the range of 0.2 < \u03bcd < 0.4. Because this is a spatial assembly we consider to show the isometric, front and top view of the folding roof structure as depicted in Fig. 10 . The desired performance, in this case, is set with a settling time in the range 1 < t s < 2 sec and a maximum overshoot of 3%. 4.3.1. Results To achieve the previously specified performance in the folding roof, let use the formula \u03c3 = 4 /t s ; thus we have \u03b1 = 2 and \u03b2 = 4 . Then, by replacing the overshoot of 3% in (11) , we have \u03b6 = 0 . 744 and \u03c8/ 2 = 0 . 736 rad. Let replace these values in the LMI conditions of Corollary 1 , the performance obtained, and a comparison with conventional stabilization approach ( Theorem 1 ) is presented in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure5.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure5.3-1.png", + "caption": "Fig. 5.3 Instances modeled in ABAQUS/explicit", + "texts": [ + " However, this present study is focused on the forming parameters and its effect of the weld line movement. In order to weld the plate of 2 mm thickness, FSW tool having 2 mm pin diameter is required. Considering that, shoulder diameter will be in the range of 4\u2013 6 mm [16]. So, the HAZ produced after welding will be in the same range. So, in present simulation study, 5 mm width of weld zone/material is considered. The tool is considered as analytical rigid body and the blank is considered as shell type deformable body (see Fig. 5.3). The blank material is considered as isotropic in nature for the ease of simulation. However anisotropic properties can be considered and accuracy of results can be further improved. The material properties used during simulations are represented in Table 5.1. The properties of the weld metal were assumed. The TWB is considered to be fabricated from AA 5754 H22 and AA 5052 H32 using the FSW process. Now onwards AA 5754 H22 and AA 5052 H32 is considered as weak base metal/weak material and strong base metal/strong material respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure48.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure48.4-1.png", + "caption": "Fig. 48.4 Image showing: a the fine-meshed resonating cavity, b extremely fine-meshed workpiece and concentrator, and c plot of element quality of meshed geometry", + "texts": [ + " So, the software tool like COMSOL Multiphysics can help in having a better understanding of the process mechanism. 48 Analysis on Thermal Characteristics of Micro-Drilled Glass \u2026 559 A 3D model of the experimental setup of microwave drilling as shown in Fig. 48.3 was developed using COMSOL Multiphysics 5.2 software to analyze the process and the behavior of workpiece while machining. The dimensions of the above model are shown in Table 48.1. 560 G. Kumar and A. K. Sharma Physics-controlled meshing is used to mesh the 3D model of the microwave drilling. Figure 48.4 shows the meshed 3D model. Fine meshing has been used to mesh the cavity (Fig. 48.4a), and extremely fine meshing has been used to mesh the concentrator and workpiece (Fig. 48.4b). The number of vertex elements, edge elements, triangular elements, and tetrahedral elements is 40, 557, 4504, and 45,401, respectively. The degrees of freedom solved for are 292,505. Figure 48.4c shows the plot of element quality of meshed geometry. The domain element statistics are as follows: average element quality = 0.653; element volume ratio = 1.924 10\u22127; mesh volume = 1.368 107; average growth rate: 2.077. The phenomenon of microwave drilling is governed by the formation of plasma at the tip of the concentrator and its subsequent interaction with the workpiece material. The formation of plasma depends upon the coupling of electromagnetic waves with the material of workpiece and concentrator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003043_jsen.2020.3007503-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003043_jsen.2020.3007503-Figure5-1.png", + "caption": "Fig. 5. Experimental platform.", + "texts": [ + " The worst condition is considered (only one joint\u2019s position can be measured by stereo vision system) to show the performance of proposed shape sensing method. The location of the measured joint will affect the performance of shape sensing method. Therefore, shape and tip position reconstruction errors, in the process of using different joint\u2019s position information, are shown in the simulations and experiments. In all the simulation and experiments, the sequence number of measured joint is unknown and it can be well estimated by proposed method. As shown in Figure 5, an experimental platform has been built firstly. It includes three parts: a wire-driven serpentine flexible manipulator, a binocular stereo vision system and a manipulator control system. The binocular stereo vision system includes two cameras, whose resolution is 5400*4800. The 3D reconstructed error of the stereo vision system is 0.18mm. The flexible manipulator with 11 vertebras is used to perform the tests. The length of each vertebra is 5mm, and the joint gap distance is 1.5mm. Each joint can rotate up to 15.31\u00b0. The vertebras are made by 3D printing, and the material is resin. The vertebra has a weight of 1g and a diameter of 9 mm. An elastic rubber tube is used to confine the joints rotations ( 22552zE I Pa kg mm = ). A pair of 0.3mm diameter steel wires is used to control the backbone bending. These parameters are used in the simulation and experiment. As shown in Figure 5, the control system includes two stepper motors which are used to pull the two cables of the flexible manipulator. And two encoders and two force sensors (Futek LSB 200) are used to measure the length change and tension force of the two cables. The precision of the encoder is 0.3\u00b0. The capacity of the force sensor is 44.5 N, and the precision of the force sensor is 0.1 N. In this section, 6 simulation cases without external payload are carried out to test the proposed shape reconstruction method" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002653_012008-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002653_012008-Figure2-1.png", + "caption": "Figure 2. SAE coordinate system", + "texts": [ + " The center of gravity point is located at the center of its y-axis, measured to be 441 mm from the front and 609 mm from the back with respect to the x-axis, and has a height of 825 mm from the ground. The vehicle has a total weight of 74 kg without any driver on it, and test was done with a driver with weight of 52 kg. ICMAA 2020 Journal of Physics: Conference Series 1519 (2020) 012008 IOP Publishing doi:10.1088/1742-6596/1519/1/012008 The coordinate system that is used is the SAE Coordinate system which goes like in figure 2 below 2.2. Geometric \u2018bicycle\u2019 modelling The basic vehicle model is by using the \u2018Bicycle\u2019 model which refers to the model made by Milliken and Milliken. This \u2018Bicycle\u2019 model represents the whole vehicle as a bicycle, where the front system is represented as a single-wheel model and so is the rear. The definition of this model includes no load transfer either on longitudinal or lateral axis, no roll, pitch, and yaw motion, constant velocity, no aerodynamic effects, no vehicle chassis and suspension compliance effects, and full position control" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002933_s10958-020-04879-x-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002933_s10958-020-04879-x-Figure2-1.png", + "caption": "Fig. 2. Scheme of a ripstik", + "texts": [ + " 148, Proceedings of the International Conference \u201cGeometric Methods in Control Theory and Mathematical Physics: Differential Equations, Integrability, and Qualitative Theory\u201d (Ryazan, September 15\u201318, 2016), 2018. 392 1072\u20133374/20/2484\u20130392 c\u00a9 2020 Springer Science+Business Media, LLC by Of and Or the points of intersection of the rotation axes lf and lr of the front and rear wheels with the inclined faces of the wedge-shaped ledges. Let Vf and Vr be the bases of the perpendiculars dropped from the points Of and Or to the longitudinal axis EfEr of the ripstik. The constant angle of slope of the wedge-shaped ledge is denoted by \u03b5 (see Fig. 2). We denote by \u03a0, \u039b, and \u03a3, respectively the supporting plane along which the ripstik moves, the inclined face of the wedge-shaped ledge, and the plane passing through the point Of parallel to the plane of the deck. The angle of slope of the front deck (i.e., the angle between the planes \u03a0 and \u03a3) is denoted by \u03b4f . The values of \u03b4f > 0 (respectively, \u03b4f < 0) correspond to the left (respectively, right) slope of the front deck, if one looks in the direction of the ripstik\u2019s motion. The slope of the rear deck is denoted by \u03b4r", + " (3) In [11], the proof of the formula (3) was based on complicated geometric constructions that greatly preclude comprehension. Here, we propose a simpler proof based on the theory of finite rotations. Let us consider the front deck of the ripstik in the position where it is not inclined. We introduce the coordinate system OXY Z whose origin O belongs to the plane \u03a0, the axis OZ is perpendicular to the plane \u03a0 and passes through the point Of , and the axis OX is parallel to the longitudinal axis EfEr of the ripstik. The unit vectors of this coordinate system are denoted by ex, ey, and ez (see Fig. 2). Let the height of the deck above the the plane \u03a0 be h. The intersection point of the longitudinal axis EfEr of the ripstik with the plane \u039b is denoted by D. Let also DOf = a. Then, with respect to the coordinate system introduced, the plane \u03a0 is defined by the equation Z = 0 and the plane \u03a3 is defined by the equation The normal vector of the plane \u039b has the form n = \u2212 sin \u03b5ex + cos \u03b5ez, (4) and the plane itself passes through the point Of whose position vector is rOf = (h\u2212 a sin \u03b5) ez. (5) Therefore, the equation of the plane \u039b is as follows: \u2212 sin \u03b5X + cos \u03b5Z = cos \u03b5 (h\u2212 a sin \u03b5) " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003963_dvm49764.2020.9243869-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003963_dvm49764.2020.9243869-Figure1-1.png", + "caption": "Fig. 1. General scheme for determining location of a defect", + "texts": [ + " First of all, it is influence of rotor vibration. Also bearings, connections, wheels, moving parts create disturbing forces. This set of disturbing forces makes the equipment to have micro movements. Generally, the equipment has vertical micro movements and micro swinging, which cause dynamic reactions on supports. These dynamic reactions can be carefully controlled and analyze may provide location of sources of disturbing. For force control strain gauges are used, that are located under the supports of the equipment (Fig. 1). Strain gauges with amplifiers and microcontroller Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 22:34:44 UTC from IEEE Xplore. Restrictions apply. allow determining dynamic reactions to the supports in real time. If it is not possible to install them, strain gauges can be glued on the supports or equipment frame. The equipment for implementation of the method contains the following: 1) strain gauge sensors located between equipment and supports or strain gauge sensors on supports; 2) amplifiers with various amplifying coefficient; 3) analog-to-digital converter; 4) microcontroller for collecting data; 5) computer for analyzing the data; 6) computer program for analyzing the data" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003674_msm49833.2020.9201736-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003674_msm49833.2020.9201736-Figure6-1.png", + "caption": "Fig. 6. Visualization of the exoskeleton within MATLAB simulation software.", + "texts": [], + "surrounding_texts": [ + "Numerical experiments of MATLAB model were conducted for one scenario - human 1.75 (m) tall and about 60 (kg) weight walking with speed of 2 (km/h). Specific limb dimensions are as followed: waist - 40 (cm), thigh - 50 (cm),calf - 40 (cm), foot - 25 (cm). Simulation model input data was prepared by filtering raw data acquired during gait identification experiment. Results of simulation in Figs. 8-10 present displacement of exoskeleton right limb joints in sagittal plane, change of velocity and acceleration in time. Authorized licensed use limited to: Carleton University. Downloaded on November 01,2020 at 22:11:02 UTC from IEEE Xplore. Restrictions apply. Ankle joint during gait reaches angular displacement in range from \u221245\u25e6 to 15\u25e6, amplitude equals 60\u25e6. Angle of knee joint reaches value in range from \u22123\u25e6 to 54\u25e6, amplitude 57\u25e6, hip joint from \u221225\u25e6 to 0\u25e6, amplitude 25\u25e6. Identifying computed signal can distinguish specific phase of gait, in time from 0 to 1.2 (s) right leg is in the stance phase, from 1.2 to 2.0 (s) in the swing phase. Signals behave periodically. The fastest change of speed in both ankle and knee joints occurs at the swing phase. Figure 11 depicts change of torque signals of specific joints in time. Rotational torque of ankle joints reach value in range from -500 (kgm2/s2) to 200 (kgm2/s2), knee joints from -800 (kgm2/s2) to 1000 (kgm2/s2), hip joint from -700 (kgm2/s2) to 950 (kgm2/s2). A significant increase of torque value happens before transition from the swing phase to the stance phase. Since identification experiment has not taken into account phenomena of the transition from the swing to the stance phase, there are differences in comparison to dynamic parameters. Figures 12 and 13 match dynamic parameters of exoskeleton respectively for identification experiment of human gait and software simulation of proposed exoskeleton. Figure 12 presesents rotational displacement of joints in time. Matching of signals\u2019 trend is high. At the first cycle of gait during the stance phase offset between both signals occurs. Authorized licensed use limited to: Carleton University. Downloaded on November 01,2020 at 22:11:02 UTC from IEEE Xplore. Restrictions apply. Figure 13 matches rotational velocity of joints respectively for identification experiment and software simulation. In comparison to chart presents angular displacement of joints, fitness of both velocity signals is poor. According to MATLAB simulation results, in the moment of transition from the swing to the stance phase velocity signal performs a rapid change of value - phase of the foot contact with the ground. Proposed concept of experiments does not provide such level of data acquisition, what will be considered at further research." + ] + }, + { + "image_filename": "designv11_80_0002774_0954406220925843-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002774_0954406220925843-Figure3-1.png", + "caption": "Figure 3. Link coordinate systems for the KUKA LBR iiwa manipulator.", + "texts": [ + " Here, it is worth noting that the first optimization criterion is the minimum index (as shown in diagram), while the second and third optimization criteria are both the maximum index. Error compensation validation and discussion for measurement configuration optimization In order to demonstrate the feasibility and universality of the proposed multilevel optimization criterion in this paper, a 7-DOF KUKA LBR iiwa serial manipulator was selected as the object of the simulation verification. Each link coordinate system was established for the KUKA as shown in Figure 3, and corresponding MD\u2013H parameters are given in Table 2. The joint angle interval of selected measurement configurations was set as 10 . The range of kinematic parameter errors were set as 0.4mm and 0.1 , respectively. Furthermore, random noise with range of 0.02 was added as a bounded error of joint disturbances; random measurement noises were defined as normal distribution variables with mean 0 and standard deviation 0.1mm, to verify the effectiveness of the multilevel optimization criterion. Finally, a 6-DOF ROKAE serial manipulator was adopted for the calibration experiment to verify the conclusion" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure5-1.png", + "caption": "Fig. 5 (A) CAD of the proposed assembly. (B) Example of an auxetic tube.", + "texts": [], + "surrounding_texts": [ + "Our primary hypothesis for a conceptualized mechanism is to vary the stiffness of our tubular surgical device by using the concept of jamming. However, unlike layer or space jamming, we use the inherent mechanism of auxetic materials, which expands upon elongation. This is more advantageous as compared to using a vacuum as there is no need for sealing, which is often one of the biggest challenges associated with jamming. By restricting the space available for the expansion of the auxetic material, it is possible to achieve jamming due to the contact force between the auxetic material and the restrictive material. A two-layered tubular system can be designed where the outer layer has fixed dimensions and is semirigid (i.e., it is only allowed to bend). The inner tube is made of an auxetic material, which expands upon elongation. However, since the available space is restricted, the inner tube will jam against the outer tube, creating friction which resists further motion. The mechanism and proposed design are shown in Figs. 4 and 5. We reasoned that we could further improve this design to increase the amount of jamming at the surface by having a counter force at the boundary pushing in the other direction. This would effectively double the normal force and, hence, double the friction. To implement this process, we used a combination of an auxetic tube material surrounded by a normal PPR material. The inner auxetic tube had a diameter of 11mm while the outer tube had a diameter of 12mm such that they fit snugly into each other with a space of 1mm as shown in Fig. 6. We theorized that when the same tensile force is applied to both the PPR and NPRmaterials, the NPRmaterial will expand while the PPR material will contract. Thus, the radius of the inner tube will increase, and the radius of the outer tube will decrease until the free space is occupied. As we continue to apply the tensile load, jamming will occur, and the resulting friction between the tubes will resist any bending forces, leading to a stiff structure. The load-bearing capacity can be increased by exerting a higher tensile force, which leads to a larger frictional force between the tubes." + ] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure7.16-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure7.16-1.png", + "caption": "Figure 7.16 Schema of membrane and the element of the boundary.", + "texts": [ + "1) with any limited right part. Eq. (7.7.1) we will examine a certain bounded space D+ of two-dimensional plane R with the smooth boundary \ud835\udee4. We will assign the boundary conditions on the border \ud835\udee4+: (i) w+(x, y) = f (x, y), (7.7.2) or (ii) \ud835\udf15w+ \ud835\udf15n (x, y) = g(x, y), (7.7.3) where the derivative of displacement along the normal to the boundary at point with the coordinates x, y is equal \ud835\udf15w \ud835\udf15n = \ud835\udf15w \ud835\udf15x (x, y)nx(x, y) + \ud835\udf15w \ud835\udf15y (x, y)ny(x, y), (7.7.4) nx, ny \u2013 are the components of the unit vector, directed along the normal (see Figure 7.16). Let us enlarge domain D+ through the boundary \ud835\udee4 to the entire plane \u2212\u221e < x, y<\u221e and continue w(x, y) and q(x, y) out of D+. Using the transformation of Fourier in the extended domain, we obtain: \u2212 1 2\ud835\udf0b \u222c \u221e \u2212\u221e \u22072w(x, y)ei(\ud835\udf09\ud835\udc65+\ud835\udf02\ud835\udc66)dxdy + \ud835\udf062W(\ud835\udf09, \ud835\udf02) = Q(\ud835\udf09, \ud835\udf02), (7.7.5) where W(\ud835\udf09, \ud835\udf02) and Q(\ud835\udf09, \ud835\udf02) \u2013 are transforms of Fourier of the desired function and load W(\ud835\udf09, \ud835\udf02) = 1 2\ud835\udf0b \u222c \u221e \u2212\u221e w(x, y)ei(\ud835\udf09\ud835\udc65+\ud835\udf02\ud835\udc66)dxdy, Q(\ud835\udf09, \ud835\udf02) = 1 2\ud835\udf0b\ud835\udc47 \u222c \u221e \u2212\u221e q(x, y)ei(\ud835\udf09\ud835\udc65+\ud835\udf02\ud835\udc66)dxdy. After taking the integral in (7.7.5) by parts, we will take into account that on the boundary \u0393 breakages of function w and its derivative along the normal n to the boundary are possible" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000614_1.5112720-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000614_1.5112720-Figure3-1.png", + "caption": "FIGURE 3: Conventional deep drawing process (a) Deep drawing test tooling-die (b). Additive manufactured insert with permeable structure (c).", + "texts": [ + " 170004-3 To put these results into perspective, in [8] the amount of lubrication required for a BMW 3 series sedan hood is about 2 g. Hence, the amount of lubrication that can be put through the specimen is sufficient for a deep drawing process. After testing the throughput rate, the required amount of lubrication during deep drawing operations can be provided by utilizing the permeable structures. FIGURE2. Schematic test set-up for the conducted throughput rate tests (a). Diagram depicting the normalized throughput rate over pressure results (b). The deep drawing process is shown in Fig. 3 a). In this study, the punch and the binder are conventionally manufactured, while the die tooling setup is as shown in Fig. 3 b). Here, a conventionally manufactured tool acts as a housing for the additively manufactured tool insert. The housing component possesses a circular channel for the lubricant transport. The lubricant is provided by the same hydraulic generator as in the previous throughput investigations. Bolted on top of the housing is the additive manufactured permeable insert, as can be seen in Fig. 3. The SLM fabricated insert, Fig. 3 c), exhibits a die radius of 50 mm and an outer diameter of 112.8 mm. 170004-4 The insert is predominantly printed in a non-permeable, dense fashion. The permeable structures are located solely in the die\u2019s flange area by utilizing a comparably increased hatch spacing of 200 \u00b5m. For purposes of testing the process, use is made of a 1 mm 1.0338 (DC04) sheet metal with a diameter of 100 mm. In order to investigate the forming process, two different tests are carried out. On the one hand, the tests are done without lubrication and on the other hand, the intrinsic lubrication is applied" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure78.8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure78.8-1.png", + "caption": "Fig. 78.8 Directional deformation of a existing and b modified FH", + "texts": [], + "surrounding_texts": [ + "The test is carried out on each sample at five different locations with three reputations. These results show that the hardness of the material is not the same throughout the hammer. At the hard-facing welded part, the hardness is high and less at the tail part. Due to this, the Fibrizer hammer frequently breaks at the welded part. The welding heat input is important welding parameter, which effects on the structure and properties of the weld metal. At location-1, sample harness result shows that the hardness varied between 110 and 124 VHN. At the location-2, Fibrizer hammer hardness is between 146 and 151 VHN, and at the location-3, hardness is between 152 and 167 VHN. 78.3.4 Cost Comparison Between Existing and Modified Fibrizer Table 78.6 indicates the cost comparison between existing and modified Fibrizer. The difference in cost between them is $11,088.00. 78.3.5 Harmonic Analysis Results of Existing and Modified Fibrizer Hammers ANSYS finite element software is used for the simulation to performing harmonic analysis. From the simulation results, total deformation, direction deformation, von Mises stresses, maximum shear stresses, and maximum amplitude are evaluated for existing and modified Fibrizer hammers. 942 T. Mathewos et al. From Figs. 78.7, 78.8, 78.9, and 78.10 analysis, results are tabulated in Table 78.7. From the results, it is understood that the total deformation in the modified Fibrizer hammer is lesser than the total deformation of the existing hammer. This shows that the modified Fibrizer hammer is more reliable than the current one. As the result shows that the maximum equivalent (von Mises) stress of current FH is increased by double, the higher the stress, the more the material will be exposed to be broken. The maximum shear stress result in the current Fibrizer hammer is more than the modified Fibrizer hammer. The smaller the radius (r) and web (h) in stepped plate of the Fibrizer hammer, the higher will be the stress. In Table 78.7, analysis results show that there is less total and directional deformation in modified Fibrizer hammer. The equivalent (von Mises) stress and max. shear stress result also much less by half from the existing. Similar manner frequency responses are verified for other surfaces also and tabulated in Table 78.8. From Table 78.8 data, the output of harmonic response has been taken in the form of amplitude which can be understood as mean stress development in any engineering components. Modified FH has more amplitude initially and gradually decreased compared to current FH. The overall behavior modified FH is taking less deformations and stresses by which life span of FH definitely will improve. 78 Design Analysis and Modification of Sugarcane Fibrizer Hammer \u2026 943 944 T. Mathewos et al." + ] + }, + { + "image_filename": "designv11_80_0003955_ecce44975.2020.9235383-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003955_ecce44975.2020.9235383-Figure2-1.png", + "caption": "Fig. 2. (a) 1 DOF and (b) 8-pole radial magnetic bearing", + "texts": [ + " 1b displays how the force density disparity may be lessened by using HV or dielectric fluid approaches. While electrostatic force density may not be staggeringly high, in the realm of flywheel energy storage, force density is a low-ranking priority. Electrostatic systems are largely 2D and can be cascaded in volumes, they also can levitate light weight materials that are inaccessible to magnetic forces. Active magnetic bearings are a relatively matured technology that are well studied and understood. A basic magnetic suspension building block is shown in Fig. 2a, which consists of two opposing magnetic actuators acting on an iron shaft. The force on the shaft from each actuator can be found by integrating the magnetic pressure term from (2) over the airgap area, leading to Fshaft = \u00b50 N2Ag 4d2 cos(22.5\u25e6)i2 (3) where N is the number of coil turns, Ag is the air-gap area, d is the air-gap distance, and i is the current through the coil. A linear force-current relationship is commonly realized using a method known as differential driving bias linearization [20], in which the coil currents are defined in terms of a fixed bias Ib and a transient control ic current. Applying the sum of the current terms to the top actuator and the difference of them to the lower actuator (see Fig. 2a) leads to the following linearized expression Fy = kiic + kyy (4) where Fy represents the difference between the forces from the upper and lower actuator, ki = \u00b50N 2AgIb d2o cos(22.5\u25e6) is the force-current gain, ky = \u00b50N 2AgI 2 b d3o cos(22.5\u25e6) is the unstable suspension stiffness and y characterizes movements about the shaft\u2019s nominal position. The equation of motion for the suspension may then be written as shown in (5). Taking the Laplace transform, and writing in canonical form gives the position transfer function of (6), where Fcy = kiic represents the controlled magnetic force on the shaft and \u03c4y = \u221a m ky . my\u0308 = kiic + kyy (5) y (s) Fcy (s) = 1/ky s2 \u03c4y \u2212 1 (6) Fig. 2b shows a practical implementation of an 8-pole radial magnetic bearing, made of two 1 DOF suspensions that are governed by (4), one for each lateral degree of freedom. At the most fundamental level, electrostatic force is created by developing an electric field (i.e., creating a potential difference) between two surfaces. In terms of an electrostatic 271 Authorized licensed use limited to: BOURNEMOUTH UNIVERSITY. Downloaded on June 20,2021 at 12:21:13 UTC from IEEE Xplore. Restrictions apply. suspension, the most basic approach to accomplish this is to apply a voltage across two electrode surfaces with the rotor object in close proximity to them, as shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001044_acc.2019.8814418-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001044_acc.2019.8814418-Figure4-1.png", + "caption": "Fig. 4. Example of a simple rectangular plate to demonstrate the correlationbased tracking algorithm. (a) Plate excited with a plane wave Einc = Ex+Ey polarized in the x- and y-directions. (b). k is the wave number of the free space.", + "texts": [ + " This method requires sufficient mesh density of the structure and a minimum frequency step size to enable correct tracking, even for higher order modes. The combination of correlation-based tracking (see Fig. 2) and frequency control (see Fig. 3) is a powerful and robust algorithm, as will be demonstrated in the next section. In this section, we investigate the correlation-based eigenvalue tracking with the proposed concept. As a generic example, a simple rectangular plate is used. Such a plate is typically used for TCM analysis of small antenna terminals [5]. The plate is excited with a plane wave polarized in the x- and y-directions [see Fig. 4(a) and (b)]. The results of correlation-based tracking algorithm for surface current correlation are shown in Figs. 5\u20137. The initial frequency step size is 50 MHz. It is to be noted that the correlation based-tracking algorithm adjusts the frequency step size in an adaptive manner, depending on the respective correlations between the modes. The tracking of the rectangular plate (see Fig. 5) shows, at the first appearance, reasonable characteristic angle curves \u03b1n , with surface current correlation", + " The disadvantage of the new algorithm is the slow convergence of the main orthogonalization in (20). The algorithm has to complete the full count of iteration steps until it recognizes the failure at the current frequency. Only after these steps, the frequency can be adjusted. For these reasons, the algorithm requires more time than the correlation-based algorithm for tracking the modes. In this section, we demonstrate the orthogonalization-based tracking algorithm used in the example shown in Fig. 4(a) and in a more complex structure, a fractal antenna. The rectangular plate is excited with a plane wave, as depicted in Fig. 4(b). The tracking of the characteristic angles \u03b1n and that of the normalized coefficient bn are shown in Figs. 11 and 12, respectively. The curves are reasonable over the entire frequency range and the effect of the degenerated modes is avoided. The tracking between modes J1 and J3, which have failed in the case of the correlation-based tracking algorithm (see Fig. 6), is now corrected. Even a simulation with a high count of frequency points yields the same result. The new recalculated characteristic modes reveal a minimized alternation with respect to the surface current distribution compared with the commonly used correlation-based tracking algorithm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000347_cleo.2019.8750415-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000347_cleo.2019.8750415-Figure1-1.png", + "caption": "Fig. 1. (a) The manufacturing procedures of the fused silica parts. (b) Fused silica glass with different 3D structures obtained from the newly developed technology. (c) The SEM images of the cross-section of the printed glass processed by different laser output power", + "texts": [ + " By combining the materials extrusion with the direct laser processing in an integrated system, a direct method for additive manufacturing of transparent fused silica glass without complicated post-heat-treatment has been demonstrated. HPMC (polymer additive) ~2-4 wt% of H2O The fused silica paste with controlled rheology was achieved from a homogeneous mixture of the pure silica powder (Purity: 99.9%, US Research Nanomaterials, Inc., TX, USA) and deionized (DI) water. The optimal composition of the fused silica paste was listed in Table 1. The general manufacturing procedures were shown in Fig. 1(a). The obtained paste was deposited layer-by-layer based on the designed CAD pattern using a precise micro-dispenser (Preeflow eco-PEN300, ViscoTec, Germany). After deposition of each layer, a high-power CO2 laser (Firestar ti100, SYNRAD Inc., WA, USA) was used to heat the paste. With laser processing, the paste can be quickly melted and fused both in plane and between the adjacent layer to form a bulk object with 3D structures. Through optimizing the laser operating parameters such as the output power, scanning speed and spot size, the transparent fused silica glass can be directly obtained without any post-heat-treatment. Fig. 1(b) showed some preliminary results of this newly developed technology. Fused silica glass with different 3D structures have been fabricated. The transparency of the laser-melted fused silica glass can be improved by tuning the laser output power. As shown in Fig. 1(c), the transparency of the laser-melted fused silica can be much improved by tuning the laser output power. Observed from the SEM images of the cross-section of the samples, the increase of the laser output power actually decreased the porosity of the laser-melted materials and thus improved the transparency. In this way, laser shows its unique potential to flexibly modify the materials properties during the manufacturing process, compared with the conventional heating method. More properties of the obtained glass such as the refractive index, light transmittance, coefficient of thermal expansion and microstructures are going to be evaluated in the next step and will be compared with the commercial fused silica glass" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002180_s00006-020-1045-1-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002180_s00006-020-1045-1-Figure3-1.png", + "caption": "Figure 3. Forces and movements of a quadrotor", + "texts": [ + " In order to apply this control law, it is necessary to compute the angular velocities \u03c9i of the motors, which are obtained from the approximation of Fi = b\u03c92 i to i-th actuator, with b the thrust factor and Fi = c\u03c92 i for the tangential force and c as the drag factor. It follows that \u23a1 \u23a2 \u23a2 \u23a3 \u03c92 1 \u03c92 2 \u03c92 3 \u03c92 4 \u23a4 \u23a5 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a2 \u23a3 0 \u2212db 0 db \u2212db 0 db \u2212db \u2212c c \u2212c c b b b b \u23a4 \u23a5 \u23a5 \u23a6 \u22121 u where d is the distance from center of mass to rotor axis, \u03c9i the i-th propeller velocity for i = {1, 2, 3, 4} as defined previously, also u = [F1, F2, F3, F4]T . Figure 3 depicts the forces and the movements of the quad-rotor. The reference for the linear position is created by using way points to following in the coordinates (xr, yr, zr) in an earth reference frame which are given by the points: (0, 0, 1), (1, 0, 1), (1, 1, 1), (0, 1, 1) and (0, 0, 1) in meters at intervals of 4 s. Table 1 shows the parameters of the quadrotor used in the simulation. To demonstrate the robustness of the proposed control law, the mass m and the inertial tensor I are increased by 30% of the initial value at t = 10 s" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000746_s40722-019-00139-6-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000746_s40722-019-00139-6-Figure6-1.png", + "caption": "Fig. 6 Sketch of WPC with RCS", + "texts": [ + " 2005) where lT and lF are the locations of the T-foil and trim tab on the ship, respectively; CLT and CLF are the lift coefficients of the T-foil and trim tab, respectively; \u03b1T is the T-foil fin angle of attack; \u03b1F is the trim tab angle of attack; AT is the area of the T-foil; AF is the area of trim tab; \u03c1 is the fluid density; and U is the ship velocity. From the experiment, the relationships between the fin angles and the T-foil/trim tab lift coefficients are shown in Fig.\u00a05. A sketch of the WPC with T-foils and trim tabs is shown in Fig.\u00a06. (3) FT = 1 2 (AT + ATF)CLT ( T , TF ) U2 FF = 1 2 AFCLF( F)U 2 MT = 1 2 lT ATCLT ( T ) U2 FF = 1 2 lF AFCLF( F)U 2, The T-foil lift coefficient, CLT, is a two-variable function of \u03b1T and \u03b1TF and can be obtained by a regression analysis (Thomas 1998) using experimental data. The stability of the WPC with RCS is herein defined as: after the WPC is disturbed by external or ride control forces, it deviates from its original equilibrium position; when the same force is canceled, the boat has the ability to return to its original equilibrium state" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001774_icmect.2019.8932157-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001774_icmect.2019.8932157-Figure3-1.png", + "caption": "Fig. 3. (a) Photograph of sun gear damage - (b) Scheme of a local tooth defect with amount of materials removed from the entire face of a tooth of sun gear.", + "texts": [ + " The shaft of WRIM is linked to a one-stage planetary gear with a fixed ring through two flexible couplings (Coupling3 and Coupling4) and a torque sensor with the frequency bandwidth of 5kHz. The input stage of the gearbox is driven by a 11kW induction motor unit with 125rpm rated speed through the main shaft, support bearings, the main bearing, and two new flexible couplings (Coupling1 and Coupling2) as show in Fig. 2.a. The tooth localized fault is produced by removing a tooth surface of sun gear with depth of 0.3mm (Fig. 3) [18]. The lumpedparameter dynamic model of the drive-train which includes the planetary gear has been established (see Fig. 2.b). This last model has been adapted to the universal dynamic lumpedparameter model with m=3 and n=6 depicted in Fig. 1.a. According to the last model of this system (see Fig. 2.b), different moments of inertia of the universal model can be obtained (for more details see [13]). The parameters of the drive train are used to adapt the universal transfer function (7) to the electromechanical system under study (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002888_j.matpr.2020.04.375-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002888_j.matpr.2020.04.375-Figure1-1.png", + "caption": "Fig. 1. Various view", + "texts": [ + " Shows the 3D modelling of Raw Material which is used for Lock ring. This design of lock ring. t and (b) Stress before heat treatment. Finite element analysis and experimental study on the deformation char.matpr.2020.04.375 Fig. 3. (a) Deformation after heat treatment and (b) Stress after heat treatment. is developed using PTC Parametric Creo 5.0 Modelling Software [6,31,36]. The diameter of the Lock ring is 4900 inch and the length of the Lock ring raw material is 3047 mm. The cross sectional area of the ring is 427 mm2. Fig. 1 shows the isometric and front view of the 3D model of EM Lock ring. Fig. 2 shows deformation and stress values before heat treatment. The total deformation of lock ring which is subjected to pressure of 7 bar & load 245166.25 N. The maximum deformation of the ring is 0.6055 mm. By the result of these analyses the ring is in safe condition. The above report shows the maximum stress acted on the ring. The maximum stress is 0.649 N/mm2. The analysis of ring is in safe condition. Fig. 3 shows deformation and stress values after heat treatment", + " In Side Gap the lock ring is placed in a gutter and checked by inserting a tool \u2018\u2018FEEDER GUAGE\u201d. The maximum allowable limit for side gap is 2.5 mm. In End Gap the side gap and the end gap is checked also by placing in the gutter itself. Using a ruler, the gap between the ends of the lock ring is measured. With respect to design calculation regarding the factor of safety and with help of analysis software it is identified that the proposed method is very much safe. The proposed method is not only better in safety but also economic. The value from Fig. 1 shows the Total Deformation of Lock ring before Heat treatment which is subjected to pressure of 7 bar & load 245166.25 N. The maximum deformation of the ring is 0. 6055 mm.The value Fig. 2 shows the Total Deformation of Lock ring after Heat treatment which is subjected Please cite this article as: P. Gurusamy, S. Hari Krishna Raj, S. Devarajan et al., acteristics of lock ring, Materials Today: Proceedings, https://doi.org/10.1016/j to pressure of 7 bar & load 245166.25 N. The maximum deformation of the ring is 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000671_012029-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000671_012029-Figure2-1.png", + "caption": "Figure 2. The parameters for successive contact points, i and j, between the chain bush and sprocket", + "texts": [ + " Series: Materials Science and Engineering 514 (2019) 012029 IOP Publishing doi:10.1088/1757-899X/514/1/012029 i iji iiji i BAOB OABAOB arccos 2 222 ; (4) from ji BOB iijiijij cosOBBBOBBBOB 2222 ; from jj BOA iij jiij j BAOA OBBAOA cos 2 222 iij jiij j BAOA OBBAOA arccos 2 222 . (5) The contact angle led to the contact point coordinates. For a known teeth number of sprocket, (z), and the wheel angular pitch (\u03c4=360/z), result the contact points coordinates (xcj, ycj). Taking account by Figure 2 and relation ij , result jA jAA jj jj sinr cosrR Cy Cx and jj jj XjYj cossin sincos T . (6) The contact point coordinates xcj \u0219i ycj result as relation jjAjjAA jjAjjAA jj jj Cj Cj cossinrsincosrR sinsinrcoscosrR Cy Cx T y x . (7) By transforming xjyj in xy( xyXjYjT ) result ii ii xyXjYj cossin sincos T . (8) Also, the relation (6) and (7) led to ijAijAA ijAijAA Cj Cj cossinrsincosrR sinsinrcoscosrR y x ", + " So, for a standardized chain transmission with unitary ratio, are considered the teeth number z=16, the bush diameter db=5.08 [mm] and also the chain pitch p= 9.525 [mm]. The transmission parameters (Figure 1a) result as follow [7]: 8248 180 . z sin p Dd mm 7443.dDD bdf mm 682 3 069050501 . d .d.R b bmax mm 5524 2 1 .R D R max f A mm 6821 .Rr maxA mm 542 2 . d r b B mm. With the previous parameters and the first method presented, there are determined the contact angle (\u03b1) depending by the contact number (i) between the chain bush and sprocket. For contact angle known values (Figure 2), the contact points coordinates (xci, yci) can be determined, when the sprocket is rotating fron 0 to the angular pitch ( ,0 ). In table 1 the different values of coordinates xci, yci (i=0...2) and \u03b1i (i=0...2), depending by rotation angle \u03b8 values, are presented. Applying the second method, for known geometrical parameters and wheel rotation angle by mounting position (\u03b8i = 0;4.5;11.25;22.5 o ) and also the angle given by the precedent chain bush ( i ), for 0F , result the contact point coordinates and also the contact angles, for the first three contacts (i=0, 1, 2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000957_icma.2019.8816405-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000957_icma.2019.8816405-Figure3-1.png", + "caption": "Fig. 3. Model of robot in Adams", + "texts": [], + "surrounding_texts": [ + "A. Virtual prototype model establishment Model our spherical robots in the SolidWorks environment. Each individual component is designed and finally assembled into an overall model. After checking the model with SolidWorks, saved in Parasolid(*.x_t) format. Open Adams, enter into the GUI and click [file/import] in the menu bar, then select File Type as Parasolid (*.xmt_txt, *.x_t.....) to import the model. The second step sets the model properties. The parts of the model in SolidWorks was represented in the Adams model, but they can\u2019t move and it looks bare. We need to change the quality, torque, color and name of each part to meet our requirements. Because there is no actual servo motor in the simulation experiment, we need to add the rotating joint to the connections of the robot, and use the rotating joint motion to move in a specified function. In addition, we also need to add ground and give contact to the part to realistically reproduce the experimental scene. The ground plane was built to improve the simulation\u2019s third dimension. We can give the rotation a simple function to see if the robot can move." + ] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure89.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure89.5-1.png", + "caption": "Fig. 89.5 Types of fissures in rice grains", + "texts": [ + " While editing, clarity was increased to full value, highlights, shadow, and whites reduced to zero value and exposure, the contrast was adjusted according to images and finally, the 89 Finite Element Analysis and Three-Point Bend Testing 1061 black and white effect was given to images. Figure 89.1 shows the images of brown rice with a different number of fissures. Figures 89.1, 89.2, 89.3, and 89.4 show the brown rice with 0, 1, 2, and more than 2 fissures, respectively. Fig. 89.2 Image of brown rice with one fissure Fig. 89.3 Image of brown rice with two fissures Fig. 89.4 Image of brown rice with more than two fissures Various types of fissures are mentioned in Figs. 89.5 and 89.6 show the image of brown rice grains (Basmati rice) with different types fissures. From Fig. 89.5, Type A fissures\u2014Fissures perpendicular to longitudinal axis of grain\u2014 Non-terminating type: The fissures present in a rice grain which are perpendicular to the longitudinal axis of grain and ended somewhere in rice grain. 89 Finite Element Analysis and Three-Point Bend Testing 1063 Type B fissures\u2014Fissures perpendicular to longitudinal axis of grain\u2014Terminating type: The fissures present in a rice grain which are perpendicular to the longitudinal axis of grain and passed throughout the rice grain along that axis" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure72.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure72.5-1.png", + "caption": "Fig. 72.5 Schematic of the boundary conditions applied", + "texts": [ + "00001 mm, which generates 20,000 elements to discritze the geometries, was chosen for the discretization of the indentation zone. The bottom surface of the indenter and the top surface of work materials formed a contact pair in the present numerical model. Surface-to-surface contact has been assigned between the contact pairs. Based on the suggestions from the literature ref. nos. [5, 6, 12], frictionless contact property is enabled between the two contact surfaces. It is because only a small portion of the indenter is considered to be in contact with the surface of the workpiece. Figure 72.5 shows the boundary conditions applied to the tool and workpiece. A fixed boundary condition is applied to the base of the work material. All the nodes on the base cannot move in any directions. The X-symmetry boundary condition is applied to the vertical surface of the workpiece as shown in Fig. 72.5. It means that the workpiece geometry and diamond tip are symmetric along with the symmetry line and both form a revolving 3D geometry around the line. Rigid body constraint is applied to the diamond indenter at the reference point. First, it is restrained in the X-direction along with all the rotational degrees of freedoms. Since the process is very slow, i.e., quasi-static loading condition is assumed. Therefore, the load applied to the indenter is the displacement/velocity load of 0.1 mm/s in the current simulation" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001670_icems.2019.8922498-Figure14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001670_icems.2019.8922498-Figure14-1.png", + "caption": "Fig. 14. Y132 Bearing Connection simulation Modal Shape", + "texts": [ + " 0 100 200 300 400 500 600 700 800 0 400 800 1200 1600 2000 R ad i a l C o n ta c t S t i ff n e s s ( P a) Numbers of Point (a) Radial Contact Stiffness Optimization Trend Authorized licensed use limited to: University of Canberra. Downloaded on July 11,2020 at 10:20:26 UTC from IEEE Xplore. Restrictions apply. The optimized bearing coupling stiffness coefficient result is 1060N/mm in the axial direction and 597N/mm in the radial direction by the multi-objective genetic optimization method. The convergence curve is shown in Fig. 13, and the first four-order mode is obtained. The mode shape is shown in Figure 14. III. MODAL VERIFICATION ON WHOLE MOTOR The above key modal parameters have been modified to ensure that the modal and measured errors of each key component of the Y series motor under the simulation calculation of a single variable are within 5%, so that the overall modal parameter is obtained.. The modified key modal parameters are imported into the Y132 model to obtain the optimized modal frequency of the whole machine, and the modal test is performed as shown as Figure 15. The test modality and the corrected modal comparison are compared in Table 12; the main modal frequency error is kept within 10%, which means there is greatly reduce in modal analysis error of the whole motor by this method above" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001788_1350650119895192-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001788_1350650119895192-Figure4-1.png", + "caption": "Figure 4. The transient coupling model.", + "texts": [ + " Therefore, the proposed mesh movement method is suitable for the transient simulation, the hexahedron mesh of the oil film can change regularly according to the relative position between the shaft, and the bearing as the figure shows. For a flexible rotor-bearing system, the system motion equations can be given as follows. M \u20acx\u00fe C _x\u00fe Kx \u00bc FU \u00fe G \u00f05\u00de where x is the general coordinate of system, M, C and K are mass, damping and stiffness matrices, respectively, FU is the oil film force and G is the gravity of rotor supported by bearing. Figure 4 shows an axial section of the journal bearing. As the bearing clearance and the axial misaligned angle is too small relative to the bearing width, the movement of the shaft and the bearing could be described with the model as shown in Figure 3. The movement distance of the shaft is defined by five points as marked in Figure 4, which are obtained by solving the rotor dynamics equations in MATLAB. For an arbitrary point P between the Pk and the Pk\u00fe1, the movement of the point P can be given according to the following interpolation. xp \u00bc xk \u00fe lk lk\u00felk\u00fe1 xk\u00fe1 xk\u00f0 \u00de yp \u00bc yk \u00fe lk lk\u00felk\u00fe1 yk\u00fe1 yk\u00f0 \u00de k \u00bc 1, 2, 3, 4 ( \u00f06\u00de On the other side, the attitude angle of the bearing, marked as ( x, y, 0), can be solved by the following equations of motions. x \u00bc z L 2 y y \u00bc z L 2 x ( \u00f07\u00de _ t\u00fe t x \u00bc _ tx \u00feMx t=Ixx _ t\u00fe t y \u00bc _ ty \u00feMy t=Iyy ( \u00f08\u00de t\u00fe t x \u00bc tx \u00fe _ x t t\u00fe t y \u00bc ty \u00fe _ y t ( \u00f09\u00de where Ixx and Iyy are the moment of inertia for the inner ring and the bearing; Mx and My are the oil film moment" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002178_978-3-030-38077-9_91-Figure18-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002178_978-3-030-38077-9_91-Figure18-1.png", + "caption": "Fig. 18. Handling performance (100 km/h) Fig. 19. Handling performance (140 km/h)", + "texts": [], + "surrounding_texts": [ + "To evaluate the transient handling response, a 4 parameters analysis was carried out by Mimuro [9]. The evaluation was carried out by giving a pulse steer input at different vehicle speeds and getting yaw rate and lateral acceleration response. For the present study, a triangular pulse of duration 0.4 s is taken as the ideal impulse is impossible to be generated in real life. The amplitude of the peak was calculated such that the lateral acceleration reached by the vehicle is less 0.4 g in order to consider the system as linear. Based on the yaw rate and lateral acceleration response, two transfer functions are obtained by using a system identification technique. The transfer functions relate the yaw rate and lateral acceleration with the steering wheel angle input. For system identification, a grey box model technique was used with the toolbox available in MATLAB\u00ae. The following equations depict the structure of transfer function and it is built based on the linear bicycle model. The TruckSim\u00ae vehicle model is nonlinear, but 1330 G. Arjunbarath et al. as an initial assumption, the same structure is used for system identification in this study. r dst \u00bc a1 1\u00fe Tf s 1\u00fe 2ns xn \u00fe s2 x2 n \u00f08a\u00de ay dst \u00bc a2 1\u00fe b1s\u00fe b2s2\u00f0 \u00de 1\u00fe 2ns xn \u00fe s2 x2 n \u00f08b\u00de The model was fitted for different velocities and the fitted model is more than 96% accurate. So, the initial assumption of the transfer function structure can be retained for further analysis. The variation in handling behavior with two static maps are studied (see Figs. 16 and 17). In both cases as the velocity increases, the area under the rhombus gets reduced. The steady state yaw rate gain was higher for Map-1 which is because of the tire operating region as mentioned in the Sect. 4.2. Overall, the handling behaviour is similar. Figures 18 and 19 compare the handling of the vehicle with Ackerman steering and with TWR (Map-1 and Map-2). Even with the redistribution of tire workload, similar handling behavior to Ackerman steering was observed. The vehicle with Map-1 is performs better than map 2 in terms of handling. In a nutshell, even with the redistribution of tire forces, the handling is similar to a vehicle with Ackerman steering. Development of Velocity Dependent Front Wheel Angle Relation 1331" + ] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure6-1.png", + "caption": "Figure 6 Brake disc deformation cloud map", + "texts": [ + " Therefore, for a differential equation expressed by an n-degree-of-freedom system, there are also n characteristic equations and n characteristic solutions, which correspond to n groups of modal shape vectors and n-order natural frequencies [11]. The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001154_978-3-030-29041-2_28-Figure9-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001154_978-3-030-29041-2_28-Figure9-1.png", + "caption": "Fig. 9. Friction coefficient during reciprocating sliding tests \u2013 lower part.", + "texts": [ + " Three possibilities were considered for material topological optimization: 35%, 50% and 75%, as depicted on Fig. 8. Functionality of the component was preserved, even for the highest mass reduction percentage, thus, allowing the geometry to be selected to proceed to production. Concerning SLM building time and costs, the material mass saved provides a competitive edge for a non-standard mould component. Despite the material savings, it\u2019s still required to keep the base of the component for cooling connections. The final aspect of the optimized component is shown on the 3D model views (Fig. 9). A hybrid building approach was also considered. This approach also provides benefits in terms of feasibility since it combines conventional manufacturing up to the component\u2019s height where geometrical complexity begins with additive manufacturing, providing all the design freedom to comply with the final application. Costs are optimized through parallel processing, less building time and less raw material for the most expensive manufacturing process. The cross-section of the hybrid nozzle bushing and its application on an injection mould are shown on Fig", + " This relationship can occur due to the wear factor. This means that with the increase of the load it is verified that the wear that occurs in the component is larger, removing a larger quantity of material from the component. Probably, this material is deposited inside the generated crater and most likely will create a tribofilm that provides better sliding conditions between the bodies in contact, which reduces the friction coefficient. Looking at the variation of the friction coefficient over time (Fig. 9), it is possible to identify an initial phase with running-in effect where this coefficient rises rather quickly, until it stabilizes in throughout the rest of the test, at the value of around 0,55. This value should be then compared to the friction coefficient of cast 316L steel under the same conditions, in order to understand how the LMD process affects this parameter. It should also be noticed that the differences of coefficient of friction observed between the upper and lower part of the component are not significant, only verifying that the measured friction coefficient at the beginning of the test was slightly higher in the case of the lower part of the specimen, however the differences are minimal and do not allow conclusions to be drawn on this aspect" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure40.5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure40.5-1.png", + "caption": "Fig. 40.5 Feed paths by vector notation method of 2D cross section and 3D model", + "texts": [ + " Birru 40 Prediction of Hot Spot and Hot \u2026 463 Hot tear was shown at the left end of the casted product, similarity index at the right side of the casted product was observed, TC-3 region the molten metal was distributed at both the sides uniformly, and TC-4 was observed as hard zone of the Al\u2013Cu alloy casted product. The solidification of molten metal has been compared with the time Vs temperature as shown in Fig. 40.4. An average of four K type thermo-couples at different regions in casting solidification process was observed. TC-1, TC-2 was hot sport and hot tear zones, respectively TC-3 is junction zone were the molten Al\u2013Cu alloy flows, and TC-4 is the hard zone. In Fig. 40.5, it was observed that the feed metal paths show the solidification direction of Al\u2013Cu alloy casting in 2D and 3D sectioned casting with feed path vector notation as mentioned below \u2022 Progressive Solidification-Cooling \u2022 Directional Solidification-Feed metal paths. The progressive solidification is represented by isothermal maps (at equal temperature). The directional solidification is represented by feed paths (with temperature gradients). 464 K. S. Pulisheru and A. K. Birru The last solidified region in casting called the hot spot which leads to volumetric shrinkage of the molten metal in the process of solidification" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003541_ccc50068.2020.9188802-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003541_ccc50068.2020.9188802-Figure1-1.png", + "caption": "Fig. 1: An example of standard formation with six UAVs. All the UAVs are evenly distributed on the circle, and the communication topology is organized in a ring form.", + "texts": [ + " Remark 1 In [15, 16], the second-order integrator model is used, which is much easier than the dynamics considered in (1). The communication network in a multi-agent system is usually modeled by an undirected graph G = {V, E}, where V = {1, 2, . . . , n} is a set of vertexes or nodes, and E \u2286 V \u00d7 V is the set of undirected edges. Let Ni denotes the set of the neighboring nodes of agent i, i.e., Ni = {j \u2208 V : (i, j) \u2208 E}. To reduce interaction among UAVs, each UAV only has two neighbors, and the communication topology is organized in a ring form, as shown in Fig. 1. Generally, the UAVs form some special shapes when executing the task, such as a line, a circle or a herringbone, etc. Inspired by the circular shape of the formation, we introduce the definition of standard formation [16], which can be seen as the fundamental formation of a multi-UAV system. Definition 1 (Standard Formation [16]) If all the UAVs are evenly distributed on the circle of radius R, and each UAV maintains a constant distance L from its neighbors, i.e., \u2016pi \u2212 pc\u2016 = R, \u2016pi \u2212 pj\u2016 = L, j \u2208 Ni where pc = [xc, yc] T denotes the center of the circle, then such a formation shape is termed standard formation. Remark 2 As indicated in [16], the control law designed for the standard formation can be easily applied for the other formation shapes by using bijective transformation, provided that any three UAVs are not in the same line in the desired formation shape. In this paper, we only target at designing control law for the standard formation. According to Definition 1, in a standard formation, the UAVs are expected to form a regular polygon. The standard formation of six UAVs is shown in Fig.1. Obviously, the relationship between the radius R and the distance L can be represented as R = L 2 sin(\u03c0n ) We use pid to represent the desired relative position of UAVi in the standard formation. To make the standard formation unique, we pre-determine the position of UAV1 on the circle as p1d = p\u2217 = [x\u2217, y\u2217]T . Thus, the desired position of UAVi can be calculated as pid = [xid, yid] T = pc + [R cos(\u03b1\u2212 (i\u2212 1)\u03b2), R sin(\u03b1\u2212 (i\u2212 1)\u03b2)]T where \u03b1 = arctan( y \u2217\u2212yc x\u2217\u2212xc ), \u03b2 = 2\u03c0 n . Choose y\u2217 = yc to simplify calculations", + " Lobs is the braking distance, which is related to the cruise speed and acceleration of the UAV. According to (13), the larger dobs, the smaller the virtual repulsion. Therefore, the UAV is subject to two forces, fi and fobs. The combined force enables the UAV to avoid crashing obstacles while keeping a standard formation. In this section, we present numerical simulations to verify the effectiveness of the proposed control law (6) and the virtual repulsion strategy. The formation control problem of six fixed-wing UAVs is considered, and the standard formation is shown in Fig. 1 with L = 40m. Thus, R = L = 40m after calculation. The parameters of the virtual leader is set as vd = 15m/s, \u03b8d = 0. The input constraints of the UAVs are set as 12m/s \u2264 vi \u2264 25m/s and |\u03c9i| \u2264 0.5rad/s. By employing the control law (6) with G = 2, the trajectories of six UAVs are shown in Fig.3. It is clear that the UAVs track the virtual leader whose path is represented by the green line. The corresponding distances between UAVs and their neighbors are shown in Fig.4(a), which implies that the distance converges to L. Besides, the corresponding distances between UAVs and the leader converge to R = 40m, as shown in Fig.4(b). Therefore, the standard formation depicted in Fig.1 is formed and kept. As for the linear and angular speed constraints of fixedwing UAVs, Fig.5(a) and Fig.5(b) show that vi and \u03c9i are bounded by vmin, vmax and \u03c9max, meaning the proposed control law is effective even with control input constraints. To demonstrate the virtual repulsion strategy, we add two obstacles in the simulations, and their parameters are shown in the Table 1. In the presence of obstacles, the position of Table 1: Parameters of Virtual Repulsion fobs Position Robs Lobs Obstacle 1 (900, 42) 10 20 Obstacle 2 (580, -6) 20 50 UAVs during maneuvering is shown in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001629_012016-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001629_012016-Figure3-1.png", + "caption": "Figure 3. 3D CAD models of test samples ((a) 0\u00ba and 90\u00ba and (b) 45\u00ba (all dimensions in mm).", + "texts": [ + " Then, the STL file was loaded to a pre-processing software called Insight, which is provided with the FDM 3D printer to be used for slicing and process parameter settings. The software, then saves the data in Chromeleon Backup Archive (CBM) file format which is ready for printing. Three printing orientations were considered: (1) along the planned loading direction (0\u00b0), (2) transverse to loading direction (90\u00b0) and (3) 45\u00b0 to the loading direction. Identical geometries were used for both 0\u00b0 and 90\u00b0 samples (Figure 3(a)) and another geometry, shown in Figure 3(b), was used for the 45\u00b0 samples. Five replicas were used for both orientations (0\u00b0 and 90\u00b0) samples and four samples were tested for 45\u00b0 samples. The samples were fabricated using the FDM machine, Fortus 450 mc. The experiment was performed on Instron 5895 using a tensile load of 250 kN. The tensile test was performed by placing the specimen into the grips of the test machine, as illustrated in Figure 4. The grips COTech IOP Conf. Series: Materials Science and Engineering 700 (2019) 012016 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002939_icredg47187.2019.191880-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002939_icredg47187.2019.191880-Figure5-1.png", + "caption": "Fig. 5. Switched reluctance generator structure (a) four-phase 8/6, (b) threephase 12/8 [16].", + "texts": [ + " The machine coils wound around the stator poles and there is no magnetic material embedded in the machine structure. The rotor also lacks any magnetic material, coil, and brush and consequently it is called cold rotor [15]. Considering the number of poles in rotor and stator, different machine arrangements emerge such as three-phase 6/4 or 12/8 and four-phase 8/6 or 16/12 in which the right and left numbers show the rotor poles and the stator poles, respectively. Examples of four-phase 8/6 and three-phase 12/8 are represented in Fig. 5. [16, 17]. In order to explain the SRG performance, it is vital to elucidate how a phase inductance changes. In a situation where the stator pole and the rotor pole are aligned, the air gap turns out to be minimum and the magnetic reluctance would then be minimum resulting in the maximum inductance. Conversely, when the poles are non-aligned, the air gap and the magnetic reluctance would be maximum whereas the inductance drops to its minimum. Clearly, the more the rotor pole approaches the stator pole, the more the inductance would be and vice versa" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003043_jsen.2020.3007503-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003043_jsen.2020.3007503-Figure3-1.png", + "caption": "Fig. 3. The stereo vision principle. rO and lO are the optical center of the right and left cameras, respectively.", + "texts": [ + " The moment balance equation of the last vertebra can be expressed: sin( ) ( ' 0.5G )cos( )n ex n ey n nM M L F F = + \u2212 + (4) where n ex n ncos( ) sin( )nxF T F T = \u2212 + , 1 1 = , 1i i i \u2212= \u2212 , n e n nsin( ) 'cos( )ny y nF T F G T = \u2212 + \u2212 , 1 1 / 2 = , 1 2T T T= + , ( )1 / 2i i i \u2212= + , 1 2( )M T T d= \u2212 and isin( / 2)iT T = . In this section, the vision-based joint\u2019s position measured method is introduced, firstly. Then the shape reconstruction method is presented. A. Vision-based Measurement Method The layout of the markers is shown in Figure 3. The markers are 3D printed on the surface of the flexible manipulator with 0.2mm location precision. On each vertebra, there are four markers with different colors (red, blue, green and black). Then stereo vision system is utilized to measure the position information of the markers. The vision-based measurement model is shown as follows [23]: 0 0 0 0 0 0 1 1 0 0 01 1 b b c b X uu f dx Y Z v f dy v Z External parameters Internal parameters = T / R T / 0 (5) where the focal length of the lens is f" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001699_1350650119893896-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001699_1350650119893896-Figure2-1.png", + "caption": "Figure 2. Coordinate graph of the journal bearing system.", + "texts": [ + " Then, the following section demonstrates the differences between the linear and nonlinear orbits at an identical working condition. The nonlinear stability boundaries are investigated in the following section. Finally, conclusions are given. The structure of the hole-entry hybrid journal bearing and the inner bearing surface are shown in Figure 1(a) and (b), respectively. The geometry is a typical hybrid journal bearing which has two rows of orifices and 12 holes in each row. The bearing aspect ratio is set as l \u00bc D=L \u00bc 1, and the location of the holes is defined as a=L \u00bc 0:25. Figure 2 is the coordinate system of the hybrid journal bearing system. Non-dimensional film thickness can be derived as19 h \u00bc 1 xJ cos zJ sin \u00f01\u00de where xJ \u00bc xJ=h0, zJ \u00bc zJ=h0, xJ, zJ\u00f0 \u00de are the shaft center coordinates. The dynamic viscosity of the lubricant is largely a function of the temperature. As a cooling system is often installed in the rotor-bearing system to take away the heat caused by the rotating shaft, the temperature can be considered unchanged. Correspondingly, the variation of the dynamic viscosity can be neglected" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001658_icems.2019.8921632-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001658_icems.2019.8921632-Figure7-1.png", + "caption": "Fig. 7. Inner air-gap radial magnetic flux density distribution of ST-MG.", + "texts": [ + " The relation between B and Az satisfies: 1 and k k k kz z r r r\u03b8\u03b8 \u2202 \u2202= = \u2212 \u2202 \u2202 A A B B (24) At the interface of kth region and (k+1)th region, the radial component of B and tangential component of H should be continuous, namely: 1 1 and k k k k k k k k r rr R r R r R r R\u03b8 \u03b8 + + = = = = = =B B H H (25) By solving the matrix equations in software MATLAB, the expressions of B can be acquired. IV. ELECTROMAGNETIC PREFORMANCE VALIDATION A FEA software tool of JMAG Designer is adopted to verify the effectiveness of proposed method. The parameters of SM-MG and ST-MG are shown in Table I. The inner air-gap and outer air-gap magnetic flux density distributions of SM-MG and ST-MG solved by FEA and HMM are shown in Fig. 4 \u2013 Fig. 7. For the SM-MG, the magnetic flux density obtained by HMM and FEA almost coincide. Notably, a little large difference can be observed for the inner air-gap magnetic flux density of ST-MG, as shown in Fig.7. That is because an equivalent method is adopted to transfer a rectangle into two sectors. Besides, a magnetic field distribution for all parts in SM-MG and ST-MG calculated via HMM and FEA are shown in Fig. 8 and Fig.9. The magnetic field distribution trends are the same for HMM results and FEA results. The strips in HMM results are caused by the assumption that the permeabilities in radial direction is a constant, which further causes error in the magnetic field distribution prediction. In addition, it can be seen that the outer air-gap magnetic flux density distribution of SM-MG is almost the same with that of ST-MG" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure22.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure22.3-1.png", + "caption": "Fig. 22.3 Coupled double ellipsoidal with three-dimensional conical heat source model", + "texts": [ + "2 shows the joint fabricated by the hybrid welding process. The thermal analysis of hybrid laser-TIG welding simulation was executed using commercially available SYSWELD software which is exclusively used for the modeling and simulation of fusion welding processes. The hybrid heat source model consisting of double ellipsoid for the arc welding and three-dimensional conical for the laser welding is used in the analysis. The geometrical parameters of the heat source models are taken from the experimental weld bead. Figure 22.3 schematically represents the combined heat source model employed in this study. The heat transfer due to confined heating while welding is mainly due to conduction mode. The amount of heat conduction relies on the materials thermophysical properties, the amount of weld metal volume produced, and weld morphology. Heat conduction defined by finding the solution for the following equation. r krT\u00f0 \u00de\u00feQ x; y; z; t\u00f0 \u00de \u00bc qCp @T=@t\u00f0 \u00de \u00f022:1\u00de where Q = Heat flux volume (W m\u22123), k = Thermal conductivity (W m\u22121 \u00b0C\u22121), q = Density (kg m\u22123), Cp = Specific heat capacity (J kg\u22121 \u00b0C\u22121), and T = Transient temperature (\u00b0C), t = time (s)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000567_itnec.2019.8729165-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000567_itnec.2019.8729165-Figure1-1.png", + "caption": "Fig. 1. Illustration of the system model", + "texts": [ + " At the beginning, each node i \u2208 V has established no links, so the nodes in swarms need to find their neighbours to build optical links. As no airborne synchronization devices are equipped, asynchronous neighbour discovery algorithm is possible. The optical head works in signal mode and randomly chooses a sector to transmit with divergence angle of \u03b8s. The interval of switching in sectors follows exponential distribution with rate of \u03bbh. We align Nh optical heads of each node side by side from 0 to 2\u03c0, then each head just able to cover a range of 2\u03c0/Nh, as shown in figure 1(a). In order to guarantee the connectivity, we evenly divided the maximum head divergence angle \u03b8h into m non-overlapping sectors, and each sector divergence angle of the beam is \u03b8s = \u03b8h/m. Each time, an optical head individually decides randomly to choose which sector to transmit. A pair of nodes can discovery each other directly if and only if their FSO beams cover the straight line of the two nodes simultaneously, as shown in Fig. 1(b). When the two nodes face each other, there is a three-way handshake method to achieve neighbour discovery. At first, the head transmits a Beacon with Hello message. When node i \u2208 V receives a Hello message from node j, it maintains alignment without changing beam direction, and sends SYN signal to node j, if node j receives SYN signal from node i, it sends ACK to node i. Specially, SYN or ACK signal contains the node information such as ID of the node itself and precise optical head pointing direction, etc", + " 2) Tw: The time when the last optical link is completed and we call it the worst situation of network setup time. We first consider the case where \u2200i \u2208 V establishes an optical link. We define Ni as the neighbours of node i, which expressed as: Ni = \u22c3 j\u2208V,j =i {j | Lij < Rd} where Lij is the Euclidean distance of node i and j, and Rd is transmission range under threshold power when \u03b8s is definite. We use node density \u03c1 to measure the nodes per unit area, then \u03c1 = N/|A|. We represent the beam coverage area of the node as S, as shown in Fig. 1(a), which meets the expression: S = Nh\u2211 i=1 Si = Nh\u03b8sR 2 d/2 (1) then the number of nodes in S follows the Poisson distribution [19], its mean is \u03c1S. We use K to describe the number of neighbour nodes in a beam sector, Prob(K) = (\u03c1S0) K \u00b7 e\u2212\u03c1S0 K! where S0 is the area of a sector. K is a random variable, we use its expectation as an estimate of number of neighbour nodes in a sector for analysis, then K = E(K) = \u03c1S0 = \u03c1\u03b8sR 2 d 2 If node j is one neighbour of node i, as the optical head has m sectors, then the rate of node i and j pointing at each other to get alignment is \u03bb0 = ( 1 m )2 \u03bbh, where ( 1 m )2 is the face-toface alignment probability of node i and j " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000609_978-3-030-12684-1_17-Figure17.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000609_978-3-030-12684-1_17-Figure17.4-1.png", + "caption": "Fig. 17.4 Lattice ABS part dimensions", + "texts": [ + " A second set of three ABS parts is printed with internal lattices. Each lattice is a 2D cell pattern printed along the build orientation, as shown in Fig. 17.3. Thus the second set of parts corresponds to the same build orientations as in Fig. 17.2, with lattices oriented vertically with respect to the base plate. The lattices are only built in the cantilever section of each part\u2014the base is left solid. The 2D cell pattern consists of 1.5 mm square holes separated by 1 mm solid walls, and surrounded by a 1 mm thick external wall. Figure 17.4 shows the outer dimensions of the lattice ABS set, which are increased from that of the solid ABS parts to allow for the internal lattice structures. A third set of parts is built in solid steel. These parts are built with the same geometry as the solid ABS parts (Fig. 17.1), except they are 3 mm thick instead of 4 mm, and have a 5 mm fillet instead of 4 mm. The three steel parts are also built in the same orientations as the three solid ABS parts (Fig. 17.2). The parts were excited by the translational oscillations of a shaker table using the Shaker Control software from Bruela & Kjaer Sound and Vibration Measurement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002982_j.mechmachtheory.2020.104001-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002982_j.mechmachtheory.2020.104001-Figure1-1.png", + "caption": "Fig. 1. Geometries of two smooth contacting convex surfaces: (a) orientation of two normals at contact, (b) frame assigned to single surface, (c) relative orientation of frames depicted in the tangent plane.", + "texts": [ + " We now briefly recapitulate the results of motion space analysis for single contact from our published work [30] . Given two smooth surfaces initially in a point contact, the interest here is to determine all possible physically admissible second-order relative motions. Since a second-order motion is completely and unambiguously described by the twist and the twist-derivative, the motion space problem boils down to finding the set of allowed twists and twist-derivatives for a given contact geometry . Consider two smooth regular surfaces that are in contact at a point ( Fig. 1 (a)). The fixed and movable surfaces are labeled S f and S m respectively and their point of contact is denoted C . The unit normal vectors to the surfaces S f and S m at the point C are n f and n m respectively, and are oppositely oriented in this work. The contact normal line is referred to as n-line (labeled n ). n is assumed to be oriented and its orientation is away from the material side of S f . The description of the local geometry of the contacting surface profiles up to second-order at the contact point requires five parameters [34] ; typically, the four principal normal curvatures belonging to the two surfaces and the relative orientation angle between the first principal normal curvature directions of the two surfaces are chosen for this purpose. The local coordinate systems x f y f z f and x m y m z m are embedded in S f and S m respectively at the common point C ( Fig. 1 (b)). A global right-handed coordinate system, XYZ , is embedded in S f at the point of contact. The positive Z -axis is away from the material side of S f and positive X -axis along the positive x f -axis. Fig. 1 (c) shows the relative orientation of the coordinate systems in the common tangent plane at the point of contact of the two surfaces. Angle \u03bc indicates the relative orientation of the surfaces S f and S m and is chosen to be the angle by which the positive x f -axis is to be rotated about the positive Z -axis to coincide with the positive x m -axis. k 1 m , k 2 m are the principal normal curvatures of S m and k 1 f , k 2 f are that of S f . \u03bc is the relative orientation angle between the first principal normal curvature directions of the two surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000542_978-3-030-20131-9_261-Figure2.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000542_978-3-030-20131-9_261-Figure2.1-1.png", + "caption": "Fig. 2.1. The over-actuated compliant joint.", + "texts": [], + "surrounding_texts": [ + "joint The internal articulated structure is designed according the \u201ccompliant mechanism\u201d concept so that the required mobility of the link (crank) (1) is obtained by means of elastic joint (fig. 1) [3]. A helical spring (2) can be taken as a compliant element and the required movements of the joint are then obtained by means of bending, expanding or contracting EAPs (3). The actuators are attached to the base (4) by their one end while by another to the crank through some circularly and symmetrically arranged levers (5), which allows using large number of actuators and control of the developed moment by adjusting the lengths of the levers. The symmetrical arrangement of actuators also simplifies calculations and makes the model more predictable [6]. Depending on the acting mode of EAPs, we can reach the necessary rotations of the joint in all three dimensions (frontal, sagittal, lateral) and one translation of the link in sagittal plane, which is needed for providing safety and regulation of the mechanism. The required capacities of actuators are determined by dynamical modeling of the compliant joint. The simplified dynamic model (damping and the inertial couplings between the links and the actuators are neglected) of the latter (Fig. 1) can be de- scribed by the following equations [7]: \ud835\udc3c\ud835\udc57(\ud835\udc5e\ud835\udc57)?\u0308?\ud835\udc57 + \ud835\udc50(\ud835\udc5e\ud835\udc57 , ?\u0307?\ud835\udc57) + \ud835\udc54(\ud835\udc5e\ud835\udc57) + \ud835\udc3e(\ud835\udc5e\ud835\udc57 \u2212 \ud835\udc5e\ud835\udc4e) = 0, (1) \ud835\udc3c\ud835\udc4e?\u0308?\ud835\udc4e + \ud835\udc3e(\ud835\udc5e\ud835\udc4e \u2212 \ud835\udc5e\ud835\udc57) = \ud835\udf0f, (2) where \ud835\udc5e\ud835\udc57 \u2208 \ud835\udc45 is the vector of joint angular positions, \ud835\udc5e\ud835\udc4e \u2208 \ud835\udc45 is the vector of actuators positions, \ud835\udc3c\ud835\udc57 is the inertia matrix of joints, \ud835\udc3c\ud835\udc4e is the inertia matrix of actuators, \ud835\udc50(\ud835\udc5e\ud835\udc57 , ?\u0307?\ud835\udc57) is the vector of Coriolis and centripetal torques, \ud835\udc54(\ud835\udc5e\ud835\udc57) is the vector of gravity torques, \ud835\udc3e is the stiffness matrix, \ud835\udf0f is the torque or vector of input signals. N. Zakaryan et al.1782 Solving (1) for \ud835\udc5e\ud835\udc5e\ud835\udc4e\ud835\udc4e and differentiating twice, we get an expression for ?\u0308?\ud835\udc5e\ud835\udc4e\ud835\udc4e. Adding (1) to (2), and inserting the expression for ?\u0308?\ud835\udc5e\ud835\udc4e\ud835\udc4e yields: \ud835\udf0f\ud835\udf0f = (\ud835\udc3c\ud835\udc3c\ud835\udc57\ud835\udc57(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57) + \ud835\udc3c\ud835\udc3c\ud835\udc4e\ud835\udc4e)?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57 + \ud835\udc50\ud835\udc50(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57) + \ud835\udc54\ud835\udc54(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57) + \ud835\udc3c\ud835\udc3c\ud835\udc4e\ud835\udc4e\ud835\udc3e\ud835\udc3e\u22121\ud835\udc37\ud835\udc37\u22121 [\ud835\udc3c\ud835\udc3c?\u0308?\ud835\udc57(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57) + +2\ud835\udc3c\ud835\udc3c?\u0307?\ud835\udc57(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57) \ud835\udc51\ud835\udc513\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc613 + \ud835\udc3c\ud835\udc3c\ud835\udc57\ud835\udc57(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57) \ud835\udc51\ud835\udc514\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc614 + ?\u0308?\ud835\udc50 (\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57, \ud835\udc51\ud835\udc513\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc613 ) + ?\u0307?\ud835\udc54(\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57)], (3) Thus the vector of input signals can be expressed as \ud835\udf0f\ud835\udf0f = \ud835\udc3c\ud835\udc3c\ud835\udc4e\ud835\udc4e (\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0307?\ud835\udc5e\ud835\udc57\ud835\udc57, ?\u0308?\ud835\udc5e\ud835\udc57\ud835\udc57, \ud835\udc51\ud835\udc513\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc613 , \ud835\udc51\ud835\udc514\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 \ud835\udc51\ud835\udc51\ud835\udc61\ud835\udc614 ). (4) Finally, the joint stiffness \ud835\udc58\ud835\udc58\ud835\udc57\ud835\udc57 is defined as the partial derivative of the torque (\ud835\udf0f\ud835\udf0f) with respect to the joint angle (\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57): \ud835\udc58\ud835\udc58\ud835\udc57\ud835\udc57 = \ud835\udf15\ud835\udf15\ud835\udf0f\ud835\udf0f \ud835\udf15\ud835\udf15\ud835\udc5e\ud835\udc5e\ud835\udc57\ud835\udc57 . As it is known, there is a dependency between the longitudinal deformation (\ud835\udf00\ud835\udf00\ud835\udc50\ud835\udc50) of a contracting/expanding EAP and the applied voltage (\ud835\udc48\ud835\udc48\ud835\udc50\ud835\udc50) [8]. \ud835\udc48\ud835\udc48\ud835\udc50\ud835\udc50 = \u00b1\u221a(\ud835\udf00\ud835\udf00\ud835\udc50\ud835\udc50 \u2212 \ud835\udf0e\ud835\udf0e\ud835\udc50\ud835\udc50 \ud835\udc38\ud835\udc38 ) \u2219 2\ud835\udc38\ud835\udc38 1 \u2212 2\ud835\udc63\ud835\udc63\ud835\udc50\ud835\udc50 , (5) where \ud835\udf08\ud835\udf08 denotes the Poisson's coefficient, \ud835\udc38\ud835\udc38 Yung module, \ud835\udc50\ud835\udc50 capacity, \ud835\udf0e\ud835\udf0e mechanical stress. It is known that electrically induced bending moment \ud835\udc40\ud835\udc40\ud835\udc4f\ud835\udc4f of the bending actuator is proportional to the input voltage \ud835\udc48\ud835\udc48\ud835\udc4f\ud835\udc4f. It is reasonable to assume that \ud835\udc40\ud835\udc40\ud835\udc4f\ud835\udc4f is proportional to the width of the sheet d [8]: \ud835\udc48\ud835\udc48\ud835\udc4f\ud835\udc4f = \ud835\udc40\ud835\udc40\ud835\udc4f\ud835\udc4f \ud835\udc3e\ud835\udc3e\ud835\udc50\ud835\udc50\ud835\udc51\ud835\udc51 , where \ud835\udc3e\ud835\udc3e\ud835\udc50\ud835\udc50 is a normalized electromechanical coupling. In summary, after determining of the required torques it will be easy to control actuators by the applied voltages, substituting (5) in (4). Necessary calculations are implemented by MSC ADAMS simulation for the joint rotations along three axes and the required characteristics of the actuators are defined. Figure 2 presents the digital model of the joint, where body 1 models the sum of the masses of the leg and other links of the mechanism, while EAPs are modeled by springs with dampers (2). Fig. 2. MSC ADAMS model of the over-actuated compliant joint. Dynamic modeling of a new over-actuated compliant joint mechanism for human\u2026 1783 Fig. 3. Force-deformation diagram for the one-actuator model during the movement in the sagittal plane (contraction/expansion). During modeling it's easy to evaluate the capacity of one actuator and then to determine it for each one analytically due to the structure symmetricity. Since EAPs are generally not so powerful, we need multiple actuators but with less power. Simple calculations show (Fig. 3) that for 100 mm of lever length we need 12 two-directional working mode actuators, if each one can develop about 444 N force (it is known that PPY-metal coil composite actuator can develop up to 500 N force). Fig. 4. Sum of the developed forces of the bending actuators. The required movement of the compliant joint can't be achieved by operating only contracting/expanding actuators, so it will be easy to calculate them in sequence: first contracting/expanding, then bending. After the first operation the link will rotate only 500, so the rest of the job will be done by bending with the appropriate actuators (Fig. 4). Force evaluation of each actuator can be done in the same manner, using again the symmetry of the structure." + ] + }, + { + "image_filename": "designv11_80_0003150_csei50228.2020.9142512-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003150_csei50228.2020.9142512-Figure7-1.png", + "caption": "Figure 7. The interface of the worm screw lift experiment", + "texts": [], + "surrounding_texts": [ + "When students take a mechanical design course, they can use the virtual remote lab to view the corresponding equipment for course preparation or review, and a step by step from easy to a complex scheme (See Figure5.) is used in building the online experiments. Some experiments interfaces are listed to better illustrate this scheme. For example, when learning drive mechanisms, the virtual lab enables students to operate fundamental experiment at the beginning, after finishing a simple one, a more complex experiment will come, this easy to difficult (see Figure6.) step by step experiments scheme in each individual chapter can consolidate students\u2019 comprehension systematically. a) The interface of the helical gear experiment Also, this step by step scheme is adopted in the whole teaching content. After the learning of each chapter, learners will have a comprehensive and global understanding of the mechanical design course, and to better strengthen the knowledge point and connect the theory of each chapter to practice, several engineering applications are also added at the final part of the course. The below pictures show the experiment interface of a worm screw lift and a truck crane, which includes worm gear and belt drive, coupling and commutator, hydraulic transmission system, the model of internal combustion engines, etc. Students can learn the characteristics of various mechanisms and have a comprehensive understanding of the principles and mechanisms through this application. 250 Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on August 17,2020 at 21:00:29 UTC from IEEE Xplore. Restrictions apply. C. Interavtive opearations In addition, interactive operations are applied to all the experiments, students can drag the left mouse button to rotate the view and drag the right mouse button to pan the view and hold the middle mouse button can zoom the whole view. Through the observation and analysis of these models, the working process of basic mechanisms and equipment becomes more interesting and vivid, which is beneficial for students to form a deeper understanding of the related mechanisms from overall angles. By combining the animations and models in the experiments, mechanical design is no longer a boring and tedious course, rather, 3D animation and interactive operation will attract students to learn and think actively. IV. CONCLUSION In this paper, the teaching process of the online mechanical design course is optimized through the systematic planning of the whole knowledge framework, and corresponding online experiments are designed. This virtual remote laboratory provides a platform where students can get complete training in the whole process of hands-on operation, mechanism and experimental result analysis, thus, students\u2019 awareness of autonomous learning and the ability to simplify and analyze problems will be cultivated. This online experiment is available for students at anywhere and anytime, thus, the contradiction between too many contents and inadequate teaching and experiment time can be solved all at once, in addition, through the interactive operation with mechanisms, mechanical design course will become more interesting and vivid. REFERENCES [1] S. Iqbal, X. Z. Zang, Y. H. Zhu, D. Hussain, J. Zhao, M. M. Gulzar and S. Rasheed, Towards moocs and their role in engineering education, 2015 7th International Conference on Information Technology in Medicine and Education (Itme) (2015), 705-709. [2] Y. Long, M. Zhang and W. F. Qiao, Survey and analysis of the application of massive open online courses (moocs) in the engineering education in china based on a survey of xuetangx, the world's largest mooc platform in the chinese language, Lect Note Netw Syst 22 (2018), 840-850. [3] 3. H. Zhixiu, S. Yongsheng, T. Xiaoying, W. Xiaojing, W. S. O. M. Chengyong, U. Tsinghua, Beijing and China, \"Research to actualize 3d animation on web-based distance learning course in engineering, 2004. [4] G. AlRegib, M. H. Hayes, E. Moore and D. B. Williams, Technology and tools to enhance distributed engineering education, P Ieee 96 (2008), no. 6, 951-969. [5] A. Alexiou, C. Bouras and E. Giannaka, Virtual laboratories in education - a cheap way for schools to obtain laboratories for all courses, by using the computer laboratory, Technology Enhanced Learning 171 (2005), 19-28. [6] R. Vuthaluru, E. Lindsay, N. Maynard, G. Ingram, M. Tade, M. Ang and H. Vuthaluru, Use of digital technologies in bridging the gap between face-to-face and remote engineering programs, 2013 10th International Conference on Remote Engineering and Virtual Instrumentation (Rev) (2013). [7] S. M. Rakshit, S. Banerjee, M. Hempel and H. Sharif, Fusion of vr and teleoperation for innovative near-presence laboratory experience in engineering education, Int Conf Electro Inf (2017), 376-381. [8] V. Potkonjak, M. Gardner, V. Callaghan, P. Mattila, C. Guetl, V. M. Petrovic and K. Jovanovic, Virtual laboratories for education in science, technology, and engineering: A review, Comput Educ 95 (2016), 309-327. 251 Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on August 17,2020 at 21:00:29 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_80_0000318_s42417-019-00117-0-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000318_s42417-019-00117-0-Figure1-1.png", + "caption": "Fig. 1 Geared rotor system with 5 shafts", + "texts": [ + " The dynamic model of the rotor system is established first, which is set up based on rotating beam elements (6-DOF at each node) and linearly meshing stiffness of engaged helical gears. Then, critical speed of the geared rotor system is calculated using the imaginary eigenvalues of the characteristic matrices. Besides, the vibration of the geared rotor system with unbalances located on different shafts is simulated by means of analytic method. Finally, vibrations of shafts are measured on a test rig to verify the simulated results. A geared rotor system with 5 shafts in integrally centrifugal compressor is described in Fig.\u00a01, and the three driven shafts, which drive impellers for work, are driven by the driving shaft through a bull gear. The model of the geared rotor system is established by finite element method, where the beam element is used to model the shafts, the supporting spring elements is used to model the supporting bearing, and the lumped mass elements is used to model the impellers and other parts with big radius. 1 3 The dynamic equations of the gear pair [17] is shown as: where Fm is the force vector; uij is the corresponding displacement vector; Kij is the mesh stiffness matrix", + " 0 400 800 1200 1600 2000 240015 20 25 30 35 40 45 f ( H z) n (r/min) Forward Backward Axial Forward Backward Coupling Axial 0 400 800 1200 1600 2000 240015 20 25 30 35 40 45 f ( H z) n (r/min) 0 800 31 32 f ( H z) n (r/min) 1200 2000 33 34 f ( H z) n (r/min) (a) rotor system without mesh coupling (b) rotor system with mesh coupling. Fig. 2 The effect of mesh coupling to the Campbell diagram 1 3 Horizontal and vertical response y and z can be expressed by where ycj , ysj , zcj , zsj \u2208 \u211d m\u00d71. After modeling the geared rotor system in Fig.\u00a01 as Eq.\u00a0(2), its nature frequencies can be calculated through Eq. (4), and the obtained Campbell diagrams of the geared rotor system are shown in Fig.\u00a02. Figure\u00a02a is the Campbell diagram without mesh coupling, where all the critical speed lines are independent from one to another. When the mesh coupling is involved, the Campbell diagram is shown in Fig.\u00a02b, where the lateral (9) y = n\u2211 j=1 ( ycj cos jt + ysj sin jt) z = n\u2211 j=1 ( zcj cos jt + zsj sin jt), 1 3 vibration, axial vibration, and the meshing vibration appear at the same time, and some critical speed lines in Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure4-1.png", + "caption": "Figure 4 Load position", + "texts": [ + " Substituting the known data into the above formula can get Nmax=18096 N. Adding load, the disc brake mainly acts on the brake disc through the friction block when the load is applied, and the load is applied to it as shown in the figure below (Figure 3). Defining constraints, apply full constraints on the center hole surface of the brake disc, apply X constraints on the inner and outer end surfaces of the disc, and apply Y and Z constraints on the friction surface of the brake disc and the friction lining. Its structure is shown in the figure below (Figure 4): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 The simulation calculation and analysis of the structural strength of the brake disc, and the results of stress solution and modal analysis deformation cloud diagram are shown in the following figure (Figure 5-6): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 From the strength calculation and cloud diagram of the brake disc structure, it can be seen that the maximum stress of the brake disc during the braking process is 52MPa, and the yield strength of the brake disc is 250MPa, so the brake disc fully meets the strength requirements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001116_ilt-05-2019-0173-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001116_ilt-05-2019-0173-Figure5-1.png", + "caption": "Figure 5 Bionic model", + "texts": [ + " For the workpieces with V-shaped, trapezoidal, circular, triangular and rectangular surfaces, the depth(h), width (s) and aspect ratio (h/s) are 0.966 0.02mm, 46 0.02mm and 0.25, respectively. In practical application, the above parameters can ensure that the required gas pressure of the air bearing is minimized when it works normally. Five kinds of bionic wing surface Bionic air bearing Siyang Gao, Jianwei Sun and Bangcheng Zhang Industrial Lubrication and Tribology Volume 72 \u00b7 Number 1 \u00b7 2020 \u00b7 122\u2013127 structures were designed to be circular using 3D software modeling, as shown in Figure 5. The experimental apparatus of Figure 6 is illustrated as follows: filter regulator is a device that adjusts the air pressure value. Ari suspension test stand is the device for experimental testing. Multimeter is a device that detects current. Balancing weight is a device for testing the mass of suspended workpieces. Battery provides electrical energy to the experimental test set. A 3D printer was used to print six kinds of suspended workpieces with a V-shaped, trapezoidal, circular arc, rectangular, smooth plane and oblique triangle surfaces" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003853_ecai50035.2020.9223138-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003853_ecai50035.2020.9223138-Figure6-1.png", + "caption": "Fig. 6. General structure of the hardware system", + "texts": [ + " 5. Simulation of the 3D printing process Authorized licensed use limited to: Middlesex University. Downloaded on November 01,2020 at 15:38:55 UTC from IEEE Xplore. Restrictions apply. In order to reduce the mass of the device and its construction costs, when making the components by 3D printing, a honeycomb structure with a filling density of 40% was chosen as a solution for the internal structures of the components [7, 8, 9]. Thus, a design model was adopted based on a standard 120 mm fan. In figure 6 you can see how the air aspirate by the fan is passed through a HEPA filter, and then captured by a turbine in order to direct it to the equipment hose, ensuring to the air a laminar flow. In figure 7, the HEPA filter is assembled with an adapter element on the fan through a bayonet type system, and the assembly consisting of the turbine, fan and HEPA filter will form a sandwich, being easy to assemble / disassemble. The entire equipment is assembled in a case specially designed to be easily assembled (figure 8), with parts for self-guiding the components, but also for self-guiding the half-cases" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003718_icra40945.2020.9197100-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003718_icra40945.2020.9197100-Figure2-1.png", + "caption": "Fig. 2: A free body diagram of each joint can be used to find the force necessary to keep the link on the surface.", + "texts": [ + " Finally, we demonstrate unique properties such as conforming to a human hand, capturing a non-adjacent object, and deploying along non-contiguous objects. Each linkage in the under-actuated gripper has two primary components: 1) a transmission capable of generating prescribed torques and 2) a series of links joined by revolute joints. In the following subsections we provide tools for the design and evaluation of this family of under-actuated grippers. Static analysis aids in the design and evaluation of these grippers. The free body diagram (FBD) is identical at each joint and can be used to generate equations for static equilibrium (Fig. 2). This analysis assumes strict stability, where each link is pushing into a surface with a positive force. These equations require a known configuration about which the transmission can be designed, referred to as the design surface [~c, ~\u03b8]. A torque at each joint \u03c4 rotates the link and must be reacted by the adjacent proximal link. Links make contact with a surface at a center of pressure (COP) at distance c and apply grip force F . Remaining engaged on the surface requires reaction forces from V x and V y (Eq", + " For a flat design surface, uniform link geometry, and uniform COP, Algorithm 1 prescribes a quadratic increase in torque from tip to base (as seen in [11]). If every link has the same COP, the rate of the quadratic governing \u03c4pat increases as the COP decreases and vice versa. Monotonic curvature actually reduces the torque required to stabilize because the shear force at each link will be divided into components and the horizontal component has a small lever arm; contributing little to the torque that must be reacted (Fig. 2). The pulley sizes can be designed with Algorithm 1, but in our case many of the pulley radii were outside our physical size constraints, so we adjusted the transmission by fitting a quadratic function between the smallest and largest practical pulley radii for both pulley series. The grip series grows whereas the retract series shrinks from tip to base (Fig. 4). This approach allows greater \u2206\u03c4 without increasing the largest pulley size. The ~\u03c4pat constructed here does not guarantee strict stability, but can be stabilized across configurations through high surface friction (Figs" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001442_042023-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001442_042023-Figure4-1.png", + "caption": "Fig 4. The schematic of strip-shaped detaching plate", + "texts": [ + " The blade mounting hole distance can meet the requirements of ordinary utility knife size and is easy to be replaced. The tool cross-section rotation radius is 55mm, which can avoid the disadvantages of the flat blade design, which is easy to be entangled. ICAMMT 2019 IOP Conf. Series: Materials Science and Engineering 631 (2019) 042023 IOP Publishing doi:10.1088/1757-899X/631/4/042023 The detachment part adopts a strip-shaped detaching plate with a length of about 6cm and a width of about 2cm. As showed in fig 4. The operating procedures of the proposed device are: Firstly, a thin insulated rope is thrown over the line that suspends the foreign objects, and the other end of the thin insulated rope is pulled to lift the device to the vicinity of the line. The device is reliably suspended on the line by the guiding action of the half-moon guiding plate. Under the action of the bearing, the device can be easily slid onto the foreign objects, and then the insulated power rope is pulled back and forth by the ground operator" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000067_9781119663546.ch1-Figure1.28-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000067_9781119663546.ch1-Figure1.28-1.png", + "caption": "Figure 1.28. The robot moves like a satellite around the Earth. For a color version of this figure, see www.iste.co.uk/jaulin/robotics.zip", + "texts": [ + " For sequentially desired directions: x\u0307d = (1, 0, 0), x\u0307d = (\"1, 0, 0), x\u0307d = (0,\"1, 0), x\u0307d = (0, 1, 0) . We get the results shown in Figure 1.27. CORRECTION FOR EXERCISE 1.15.\u2013 (Follow the equator) 1) The rotation matrices to go from one frame to another are Rij = RT iRj . We have R0 = I3 R1 = Reuler - 0, 0, '\u0307Et . R2 = Reuler (0, 0, atan2 (p2, p1)) \u00b7RT euler - 0,\"\" 2 + asin p3 &p& ,\" \" 2 . 2) The kinematic equations, taken from [1.13], are: ( ) * p\u0307 = R3 \u00b7 v3 R\u03073 = R3 \u00b7 (!3') v\u03073 = a3 +RT 3 \u00b7 g (p)\" !3 ' v3 On the simulation (see Figure 1.28), we observe that the trajectory corresponds to an ellipse which is consistent with the behavior of a satellite. The rotation of the body is due to the initial conditions. 3) The dynamic model is composed of the kinematic model, to which we add the following state equation to generate the inputs a3,!3 (see Figure 1.29): ( ) * !\u03073 = \"!3 +RT 3 \u00b7 !E + u(3 a3 = RT 3 \u00b7 (!E ' p)\" v3 +RT 3 (!E ' (!E ' p)\" g (p)) + ua3 This dynamic (left) block has !3 among the state variables. Let us explain the first equation: " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002341_b978-0-12-812162-7.00004-7-Figure4.10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002341_b978-0-12-812162-7.00004-7-Figure4.10-1.png", + "caption": "FIGURE 4.10 Type of movements (A) rolling, (B) spinning, and (C) sliding.", + "texts": [ + " Arthrokinematics is the specific movements that occur at the articulating joint surfaces are referred to as arthrokinematic movement (Lippert, 2011; Floyd, 2015). These are the motions of bony segments as well as the motion of the joint surfaces in relation to another. One of the joint surfaces is relatively stable and serves as a base for the motion, whereas the other surface moves on this fixed base. When one joint surface moves relative to the other, spin, roll, slide or combinations occur. The terms rolling, spinning, and sliding is used to describe the type of motion that the moving part performs (Kaufman and An, 2008) (Fig. 4.10). Rolling refers a series of points on one articulating surface come into contact with series of points on another articulating surface. Rolling occurs in the direction of movement. An example of rolling is femoral condyles rolling on fixed tibial plateau during flexion and extension in standing. Spinning refers to a rotational movement around a longitudinal axis. The same point then on a moving surface creates an arc of a circle as the bone spins. An example of spinning is the radial head of the proximal humeroradial joint during supination and pronation of the forearm" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002982_j.mechmachtheory.2020.104001-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002982_j.mechmachtheory.2020.104001-Figure10-1.png", + "caption": "Fig. 10. Proposed pair in SFC using two symmetric special contacts. (a) shapes in contact, (b) Dupin indicatrices of the surfaces at the two points of contact.", + "texts": [ + "5 that a revolute pair can be obtained using two concave contacts with the axis of rotation being the common n-line of the contacts. If somehow the spin about this n-line is restrained, then that pair will be in SFC. To achieve this, it is proposed to use special contact geometry mentioned in section 3.3.2 of [30] . Unlike a general contact, in a special contact geometry, the normal curvatures of the two surfaces S m and S f match along a tangential direction. This is also the direction of the location of point of tangency of the two Dupin indicatrices (see Fig. 10 in [30] ). The analysis in section 5.2.1 of [30] shows that the characteristic of M (t) at a contact, i.e., either penetration or separation or persistence of contact, is based upon the sign of B \u2032\u2032 ( Eq. (A.1) in Appendix 3 ). If B \u2032\u2032 < 0, then irrespective of the sign of A there shall be penetration. On the other hand, if B \u2032\u2032 > 0, then the further analysis is similar to that of the general contact and sign of A (see Eq. (1) above) comes into picture. If B \u2032\u2032 = 0 , then the signs of A \u2032\u2032 ( Eq. (55) in [30] ) and A decide the characteristic of motion at a contact. In Fig. 10 , two objects are in contact over two special contacting geometries. Suppose a M (t) having its associated twist coordinate N a = 0 be impressed upon the movable object. For the analysis at contact- a , the sign of B \u2032\u2032 a dependent upon the sign of N a (see Eq. (A.1) in Appendix 3 ). In the coordinate system at contact- b , the corresponding spin component of the angular velocity along Z b is labeled N b and we have N b = \u2212N a . Hence, for similar contacting geometries at the two contacts, the signs of B \u2032\u2032 a and B \u2032\u2032 b are always opposite as calculated using Eq. (A.1) in Appendix 3 . This means that for any arbitrary N a , one of B \u2032\u2032 a and B \u2032\u2032 b is strictly negative; this implies penetration at one of the contacts and such a M (t) is restrained. This can also be understood from the Dupin indicatrices shown in Fig. 10 (b). For a CCW spin about Z a , Dup ( S ma ) is rotated CW about Z a , then Dup ( S mb ) is rotated CCW about Z b and intersects Dup ( S fb ); this is not admissible for a proper contact. Thus M (t) with N a > 0 is restrained by contact- b . In a similar fashion, it can be argued that M (t) having N a < 0 will be restrained by contact- a . Now coming to case when M (t) has the associated twist coordinate N a = 0 , then we have B \u2032\u2032 a = 0 and B \u2032\u2032 b = 0 . The analysis requires the signs of A \u2032\u2032 a and A \u2032\u2032 b (to be computed using Eq. (A.2) in Appendix 3 ) and also signs of A a and A b to determine the characteristics of the motion at the two contacts. The algebraic expressions for A \u2032\u2032 a , A \u2032\u2032 b , A a and A b for these special geometries are mathematically involved and do not yield to the algebraic manipulation techniques used in the above examples to establish SFC. This problem is under investigation and it is believed that two symmetric special contact geometries as shown in the Fig. 10 are sufficient for SFC . 6. Discussion Consider two objects in multiple point contacts, say one is fixed and the other movable. If we perform kinematic inversion and exchange the roles of fixed and movable objects while keeping the geometries same, the relative motion space remains the same. In the examples of SFC and KP, as a result of kinematic inversion, we can obtain pairs with altered shapes on the contacting objects but retaining the same kinematic characteristics. In the examples of SFC and KP, for the limiting case of spheres reducing to points, we have k f \u2192 + \u221e " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003730_tmag.2020.3027291-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003730_tmag.2020.3027291-Figure1-1.png", + "caption": "Fig. 1. The basic model of PMTM", + "texts": [ + "ieee.org. Digital Object Identifier 10.1109/TMAG.2019.2906084 Authorized licensed use limited to: Auckland University of Technology. Downloaded on October 04,2020 at 08:03:37 UTC from IEEE Xplore. Restrictions apply. 0018-9464 (c) 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. The basic structure of the permanent magnet toroidal motor is shown in Fig. 1. It is mainly composed of (1) central worm stator, (2) planet carrier rotor, (3) planet gears and (4) outer stator. The central worm stator is made of silicon steel sheet and its surface is provided with uniform spiral winding slots, which are used to place the three-phase armature windings to generate the rotating spiral magnetic field. The planet gears are made of soft magnetic material, and it drives the planet carrier rotor to rotate. The permanent magnet teeth are evenly distributed along the circumference of the planet gear" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002206_00207179.2020.1734237-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002206_00207179.2020.1734237-Figure1-1.png", + "caption": "Figure 1. Open kinematic chain.", + "texts": [ + " In this section, we describe the underwater vehicle kinematic model using screws (Davidson &Hunt, 2004; Hunt, 1987) combined with the Davies\u2019 method (Davies, 1981) to represent its differential motion. This representation takes to account a coordinate transformation between Euler angles andOctonions, and the virtual kinematic chain associated with the vehicle movement, and consists of the main contributions about kinematic modelling of such vehicles. We describe the vehicle movement by an open kinematic chain movement. We define it as open kinematic chain as a set of links organised in series by n joints where the relativemovement happens between the links (Figure 1). Suppose $i i = 1, . . . , n screws that describe the relative movement between the links, i.e. $1 describes the joint movement 1 (among the link A and the base), $2 describes the joint movement 2 (among the link B and the link A), $3 describes the joint movement 3 (among the link C and the link B) and so on. Figure 2. Kinematic chain that represents the underwater vehicle. In this case, themovement of the last link from the chain (U) in relation to the base is given by $U = n\u2211 i=1 $i. (1) Each screw can be represented by its magnitude and by its normalised twist $\u0302, which results in $U = n\u2211 i=1 $\u0302i i" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003824_aiea51086.2020.00073-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003824_aiea51086.2020.00073-Figure4-1.png", + "caption": "Figure 4 Rear view of the clamping mechanism", + "texts": [ + " This often requires the preparation of multiple installation tools in the actual connection work In order to cope with the J-type clamps of different sizes and models, it brings a lot of inconvenience to the relevant operators, and also increases the design difficulty of the clamp installation tool[8]. III. STRUCTURE AND WORKING PRINCIPLE OF THE DEVICE The clamping device for installing tributary wires is a nonfixed position wire clamping method, with a large contact area and a fixed contact position, the clamping force is adjustable, and can be well adapted to different types of J components and different Diameter wires. The structure is shown in Figure 3 and Figure 4, including: Branch cable clamping arm, stop crossbar, fixed crossbar, moving crossbar, turning arm, rotating handle, connector, extension spring and other parts; 1- rotating handle; 2- turning arm; 3- stop crossbar; 4- branch cable clamping arm; 5- extension spring; 6- fixed crossbar; 7- moving crossbar; 8- connector The bottom end of the branch cable clamping arm is hinged to the shunt device, and the entire clamping arm can swing at a certain angle around the hinge point. The clamping arm is used to clamp the branch cable, so that the branch cable can be temporarily fixed in the lead groove of the J-type clamp" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001288_j.matpr.2019.08.229-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001288_j.matpr.2019.08.229-Figure2-1.png", + "caption": "Fig. 2. Illustration of a) Laminated and,", + "texts": [ + " In particle reinforced composites, the particles of different sizes and shapes are randomly dispersed in matrix. It results in increase the modulus of matrix. Coming to fiber-reinforced composites, the properties mainly depends on fiber properties and the applied load. Based on fiber arrangement they may be continuous-or discontinuous-aligned composites shown in Fig. 1. In structural composites, properties are mainly depends on design and reinforcement. The common structural composites are shown in Fig. 2(a) & (b). Tabakov and Summers [1] are used stress function method, to determine circumferential, radial and tangential stresses, by considering cylindrical cylinder closed at both ends. Khan [2] have done the stress distribution analysis for a horizontal pressure cylinder with saddle supports. Lei Zu et al. [3] presented a simple methodology with unequal polar openings. Alibeigloo [4] analyzed the free and static vibration features by using differential quadrature methods. Hocine et al. [5] presents analytical and experimental analysis of cylindrical cylinders filled with hydrogen" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002504_9781119592945-Figure3.7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002504_9781119592945-Figure3.7-1.png", + "caption": "Figure 3.7 (a) Form of optimally shifted bar with \ud835\udefdl = 0.86; and (b) form of oscillations of a bar of an optimal configuration (1), and constant (2) cross section.", + "texts": [ + "935 from the mass of the bar with constant cross-section. When \ud835\udefdl is equal to 0.65 and 1.08, which corresponds to M0/M equaling to 0.5 and 2, then the savings of mass will be 0.97 and 0.845, respectively. It turns out that the savings in the mass of a construction is more with greater values of \ud835\udefdl, which corresponds to a case when the mass of a bar itself is greater than the concentrated mass. This is understandable, since the value of the concentrated mass is fixed and the minimum of weight is obtained by the variation of the mass of bar. Figure 3.7 shows the optimum configuration of a bar with \ud835\udefdl = 0.86 and the modes of free oscillations of the fundamental tone of optimum and constant cross-sections of bars. The configuration of the bar depends on the value of \ud835\udefdl, and the greater are the values of \ud835\udefdl the more pointed (peaked) it becomes. 3.6 The Oscillations of Flexural-Shifted (Bending-Shifted) Bars Under the Seismic Impacts Extensive literature has been dedicated to the oscillation of bent beams, taking into account the shearing strain and rotary inertia (or the so-called beams of Timoshenko) [1, 146, and 388]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000956_pc.2019.8815034-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000956_pc.2019.8815034-Figure1-1.png", + "caption": "Fig. 1: (a) laboratory model and (b) a simplified schematic of the inverted pendulum.", + "texts": [ + " Section II briefly describes modeling of the inverted pendulum-cart system and formulation of the continuous-discrete extended Kalman filter. Section III describes controller design exploiting an energyefficient swing-up and MPC-based stabilization of the system. Section IV next presents the utilized laboratory model of the system and the hardware\u2013software setup followed by obtained experimental results and their discussion. The paper concludes with some final remarks and objectives of future work. 978-1-7281-3758-2/19/$31.00 c\u00a92019 IEEE 209 The target laboratory model to be controlled is depicted in Fig. 1a. The pendulum is mounted on a free rotary joint that is fixed to the cart that can move only in the horizontal direction. The cart can slide on the rail, and it is driven by the stepper motor via toothed belt. The position of the cart and the angle of the pendulum are measured with the incremental encoders. A mathematical model of this inverted pendulum, schematically illustrated in Fig. 1b, can be described by the following equations of motion derived by the second Newton\u2019s law or the Euler-Lagrange equation: (m1 +m2) q\u0308 +m2l\u03b8\u0308 cos \u03b8 \u2212m2l\u03b8\u0307 2 sin \u03b8 = F, (1a) l\u03b8\u0308 + q\u0308 cos \u03b8 \u2212 g sin \u03b8 = \u2212b\u03b8 \u03b8\u0307, (1b) where m1 and m2 represent the mass of cart and pendulum, respectively, g is the acceleration due to gravity, l denotes the distance from pivot to pendulum\u2019s center of gravity and b\u03b8 is the damping coefficient for pendulum. q, q\u0307, q\u0308 denote position, velocity and acceleration of the cart, \u03b8, \u03b8\u0307, \u03b8\u0308 denote angular displacement, velocity and acceleration of the pendulum, and F is the external force acting on the cart", + " Before proceeding with the controller design, we must first design an appropriate nonlinear observer which would enable to reconstruct the state vector x(k) from the (possibly noisy) position measurements y(k). 1This is due to the fact that the actual stepper motor can start, stop or reverse \u201cinstantly\u201d. Therefore, the reaction of the pendulum arm on the cart can be neglected. This interesting observation was pointed out and experimentally validated on the identical laboratory system in [7]. 2In reality, in case of the laboratory model in Fig. 1a zero position of the pendulum is its downward stable equilibrium (where the encoder\u2019s count is initiated), which is however recalculated so as to match the upward unstable equilibrium for the sake of the stabilizing controller. In order to provide reliable state estimates for both swingup and stabilizing controller we need to consider the full nonlinear model of the system, (2), which is valid over the entire feasible operating space of the cart and the pendulum. To this end we propose to employ an extended Kalman filter (EKF)\u2014 the nonlinear version of the well-known Kalman filter which at each sampling instant linearizes the nonlinear model about the current operating point using Jacobi linearization. EKF is fit for our purpose, as the system\u2019s state-space model (2a)\u2013(2b) is continuously differentiable and has only mild nonlinearities. As most physical systems, the laboratory inverted pendulum system in Fig. 1a is represented as a continuous-time model while discrete-time measurements are taken for state estimation via a digital processor. Therefore, to design the so-called continuous-discrete EKF [8, Table 6.1-1] let us assume that the system model and the measurement model are given by: x\u0307(t) = f(x(t), u(t)) +w(t), (5a) yk = h(xk) + vk, (5b) where w(t) and vk are additive process and observation noises which are both assumed to be zero-mean multivariate Gaussian noises with covariance Q(t) and Rk, respectively", + " Aqpu \u2264 b\u0303qp(x\u2212x r) , (12b) where the control error dependent vectors g\u0303 and b\u0303qp have been slightly modified to enable reference tracking in terms of cart position and to shift state constraints adequately. The overall control strategy of the inverted pendulum system eventually takes form of the following switching control law4: u= satu,u(usu(x\u0302)) , if x\u03023 /\u2208 [\u2212\u03b8s, \u03b8s], u\u22c6 0\u2190 min u 1 2 uTHu+uTg\u0303(x\u0302\u2212xr) s.t. Aqpu \u2264 b\u0303qp(x\u0302\u2212x r) , if x\u03023\u2208 [\u2212\u03b8s, \u03b8s], (13) where \u00b1\u03b8s is pendulum angle at which the controllers switch from swing-up5 per (9) to balancing-tracking per (12). The proposed scheme to control the laboratory device shown in Fig. 1a and described in Sect. II was implemented using the MathWorks\u2019 Simulink Real-Time prototyping suite on a target PC equipped with 2.4 GHz CPU, 4 GB of RAM and a National Instruments PCI-6601 counter/timer device providing the necessary input/output interface. 4Note that the actual implementation uses the state estimates x\u0302. 5Note that in (13) the input usu is saturated at the same bounds as used by the stabilizing controller, according to satu,u(usu)\u2261max (u,min (u, usu)). Table I lists the values of parameters of system model, state observer and controller as used in the laboratory experiments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure12-1.png", + "caption": "Figure 12. Ten-bar mechanism.", + "texts": [], + "surrounding_texts": [ + "The primary objective of this study was to identify the joint that producesthe highest error in a mechanism due to clearance. Simulations were carried out for all mechanisms described in the previous section by introducing clearances in different joints, one at a time. The effect of different speeds (of crank) and clearance sizes was also studied. Some important results are discussed here. The results for the four bar mechanisms have been summarized in figures 13, 14 and 15. These show that Joint 3 is the most critical joint and the values vary significantly for 0.02 mm but they are very similar for 0.5 mm. So, sensitivity to speed is high for low clearance. Figures 16, 17 and 18 show mean deviation at different joints of different inversions of the Watt chain at 0.1 mm radial clearance. The trend of all the curves is the same. Joint 3 is the most critical joint in all the three mechanisms followed by joints 2, 4 and 5, i.e., the criticality order here is: Joint 3[ Joint 2[Joint 4[ Joint 5. Joint 6 in figures 5 and 6 is the grounded joint which is least critical of them all. Figures 19 and 20 show the criticality graphs for Stephenson1 mechanism and Stephenson1 mechanism with slider, respectively. Both the graphs show the same trend where the deviation is maximum at the output link and decreases as one proceeds to the other end of the mechanism. A common trend which seems to emerge is that the first joint next to the crank, which has a link connected to ground, is the most critical joint. After that the criticality ranking goes towards the crank and then goes further down to the joints in order of their distances from the crank. We would examine the validity of this hypothesis in the rest of the study. Figures 21 and 22 show the criticality graphs for Stephenson2 mechanism (1) and Stephenson2 mechanism (2). The order of criticality for both the mechanisms is: Joint4 [ Joint3 [ Joint2 [ Joint5. This is as per the hypothesis stated earlier. The most critical joint is the joint next to the crank, which has a link connected to the ground (i.e., Joint 4). The next most critical joints are the ones towards the crank, i.e., Joint 3 and Joint 2. Then comes the joints which are left, i.e., Joint 5. Joint 6 is the grounded joint. The two graphs look very similar as the structures of mechanisms are the same. The deviations increase significantly with increase in speed, if the clearance is at Joint 4. The values in first graph are slightly higher than the corresponding values in the second graph. The overall conclusion is that deviations are more, if the output link is closer to the crank. Similar results are seen in eight and ten-bar mechanisms as shown in figures 23 and 24. The joint next to the crank is the most critical joint. It should also be noted that the grounded joints also show a trend. The grounded joint in the first loop has a significant effect but rest of the grounded joints have very little effect on the output. Error in output generally increases with increase in the crank speed (expect for an aberration shown in figure 19). This happens for all clearance sizes. However, the change in mean deviation is quite small even with a 5 times increase in speed from 100 to 500 rpm. Figures 25 and 26 show the variation of mean deviation with increase in speed when a clearance of 0.1 mm is at Joint 2 and Joint 3, respectively. It can be noticed that the deviations usually increase with speed. This means that a machine working at higher speeds is expected to produce more error than a machine working at lower speeds. Increase in radial clearance size brings an almost linear increase in the deviation values as seen in figures 27 and 28. It is also worth noting from figure 27 here that the deviation of slider is nearly 50% for 0.1 mm clearance (0.0501 is 50% of 0.1) and it goes on increasing to 80% for 0.5 mm clearance (0.3977 is 80% of 0.5). This happens because at lower values of clearance, the number of collisions between journal and bearing is more. Due to more number of collisions, the journal is sent back to the center again and again, and thus it remains at the center for a larger amount of time as compared to the case of large clearances. Thus, in larger clearances, the journal appears to remain closer to the walls of the bearing for an extended period and spends less time going to the diametrically opposite wall." + ] + }, + { + "image_filename": "designv11_80_0001661_icems.2019.8921927-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001661_icems.2019.8921927-Figure7-1.png", + "caption": "Fig. 7. Total deformation nephogram. (a) non-uniform. (b) uniform.", + "texts": [ + " 3) The rotor core is assumed as an entity and the lamination effect on material properties is negligible. The International Electro Technical Commission standard IEC60034-1 stipulates that the motor should be able to withstand 1.2 times the maximum speed. Therefore, the mechanical state under the rotor with a rotational speed of 14400rmp is taken as the research object when performing stress simulation. The mechanical stress simulation nephograms and total deformation nephogram of the two rotors are shown in Fig. 6 and Fig. 7. The simulation results show that the maximum mechanical stress (MMS) of rotor with uniform air gap is 456.68 MPa, a little smaller than the rotor with non-uniform air gap , which is 464.27 MPa. The MMS point of both rotors on bilateral bridge. In this paper, the structure of the outer ring of the rotor does not affect the position of the MMS point and the MMS value. Moreover, the total deformation of the two rotors is basically the same. IV. CONCLUSION The simulation results of one particular design are demonstrated in this paper" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002032_3352593.3352615-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002032_3352593.3352615-Figure3-1.png", + "caption": "Figure 3: Rover frame description with respect to global frame", + "texts": [ + " , 6 is set at each rover wheel at the contact point with the ground, where z-axis is normal to the terrain surface and the x-axis at that point, while yaxis is parallel to the sidewise direction of wheel. Fig. 2 shows angle of z-axis of Ci with respect to the z-axis of wheel-axle frame Ai .This angle is called the contact angle (\u03b4i). On a flat surface, the z-axis of the wheel-contact frame and the wheel-axle frame are always coinciding (\u03b4i = 0), but contact angle may take some nonzero value on rough terrains. Our assumption here is that the contact position of the first wheel (C1) is given in rover reference coordinate frame R. Fig.3 shows the different frame attached with the 10 DOF rover and wireframe structure of rocker-bogie. The rover base configuration in vector form is denoted as \ud835\udc3b = [\ud835\udc4b, \ud835\udc4c, \ud835\udc4d, \ud835\udf19\ud835\udc65 , \ud835\udf19\ud835\udc66 , \ud835\udf19\ud835\udc67] \ud835\udc47 which is defined with respect to the global coordinate frame G, where [\ud835\udc4b, \ud835\udc4c, \ud835\udc4d] is the position of the origin of R, orientation angles [\ud835\udf19\ud835\udc65 , \ud835\udf19\ud835\udc66 , \ud835\udf19\ud835\udc67], heading angle \ud835\udf19\ud835\udc67, pitch \ud835\udf19\ud835\udc66, and roll \ud835\udf19\ud835\udc65 as shown in Fig. 4. Using the DH table the following transformation matrices can be obtained \ud835\udc47\ud835\udc37 \ud835\udc45; \ud835\udc47\ud835\udc35\ud835\udc56 \ud835\udc37 , \ud835\udc56 = 1,2; \ud835\udc47\ud835\udc46\ud835\udc57 \ud835\udc35\ud835\udc56 , j = 1,2,3,4; \ud835\udc47\ud835\udc46\ud835\udc58 \ud835\udc37 , k=5,6, where Bi denotes the bogie frames, D differential frame , and Sj, and Sk denote auxiliary steering frames (attached to rocker arm)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003347_s11668-020-00975-x-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003347_s11668-020-00975-x-Figure1-1.png", + "caption": "Fig. 1 The waviness error on the planet bearing (M and x are the external moment and input speed)", + "texts": [ + " [28] Fcj \u00bc 2 Z l 0 2:447g0xR Dwx=2\u00f0 \u00de2 ffiffiffiffi h0 p dx \u00f0Eq 6\u00de where g0 is the oil viscosity; l is roller length; x is the roller length direction; and h0 is the oil film thickness, which is given as Harris and Kotzalas [27], Bernard [13] and Schaeffler [25] h0 \u00bc 1:95 a0g0u\u00f0 \u00de8=11R 4=11 e E 1=11 0 q1=11 \u00f0Eq 7\u00de where u is the relative velocity, a0 is the pressure viscosity factor, Re is the equivalent radius, E0 is the equivalent elastic modulus, and q is the force per unit length. The viscosity moment Md is given as Md \u00bc XZ j\u00bc1 1 8 Cdq0Dwl dmxm\u00f0 \u00de2 dm 2 \u00f0Eq 8\u00de where Cd is the drag factor, which is 3.0; and q0 is the oil density. Furthermore, the total power loss of the planet bearing is given by P \u00bc 1:05 10 4Mbnb: \u00f0Eq 9\u00de Calculation Method for the Abnormal Bearing When the waviness error generates on the bearing bodies, the contact forces in the bearing should be changed [22] as shown in Fig. 1. Thus, it can change the friction moment and power loss of the planet bearing. The waviness on the inner race is given by Lynagh et al. [29] pij \u00bc Xn s\u00bc1 Ais cos s xi xc\u00f0 \u00det \u00fe 2ps j 1\u00f0 \u00de Z \u00fe ais \u00f0Eq 10\u00de where s is the waviness order; Ais is the inner race waviness amplitude; xc is the cage speed; t is the time; and ais is the initial angular angle for inner race waviness. Moreover, the waviness on the outer race is written as poj \u00bc Xn s\u00bc1 Aos cos s xo xc\u00f0 \u00det \u00fe 2ps j 1\u00f0 \u00de Z \u00fe aos \u00f0Eq 11\u00de where Aos is the outer race waviness amplitude; xo is the outer race speed; and aos is the initial angular angle for outer race waviness", + " The changes of the contact forces in the planet bearing due to the waviness error can be obtained by using the Taylor series, which is given as Qw \u00bc knd 10 9 \u00fe 10 9 knd 1 9Dd\u00fe 5 81 knd 8 9 Dd\u00f0 \u00de2 \u00f0Eq 12\u00de where Dd is the deflection changes caused by the waviness error, which can be defined as the amplitude of the waviness error as given in Eqs 10 and 11. The deflection change (Dd) is not the only cause of the power loss. To consider the effect of the waviness error on the power loss of the planet bearing, Eq 12 could be submitted into Eqs 1\u20139. The used geometric parameters of the planet bearing are shown in Table 1, which are from Talbot et al. [19]. Model Validation Figure 1 plots the power loss from the experimental data in Talbot et al. [19] and the proposed method under different speed conditions. Here, the oil temperature is 90 C, and the input speed is 1800 and 3500 r/min, respectively. In Fig. 1, the power loss of the planet bearing from the proposed method and experimental data in Talbot et al. [19] increases with the moment. Note that the results from the two methods are very similar. As shown in Fig. 2, the average errors between the experimental data and the data from the proposed model are 3.8 and 4.4%, respectively. It can give some validation for the proposed method. Effects of Waviness Error and Pinion Bore Diameter on the Power Loss Figure 3 shows the effects of the waviness error and pinion bore diameter on the power loss of the planet bearing" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002956_022027-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002956_022027-Figure2-1.png", + "caption": "Figure 2. Calibration principle.", + "texts": [ + " The transformation matrix between the tool coordinate system and the base coordinate system of robot can be expressed as: base 0 6 tool 6 tool*T T T .If only the position relation of the tool coordinate is considered, the transformation relation between the end center of the robot tool and the base coordinate system is expressed as follows: base 0 6 0 tool 6 tool 6*P N P P Controlling TCP to contacts a fixed point from different directions to ensure that the TCP position was fixed when computer collects data, as shown in Figure 2. In this case, different calibration points have different joint angles because of different orientations, but the actual position of TCP is unchanged. IWAACE 2020 Journal of Physics: Conference Series 1550 (2020) 022027 IOP Publishing doi:10.1088/1742-6596/1550/2/022027 The position error between two calibration points is expressed by 2 1 1- - 0n n...P P P P P . However, there is a deviation in the TCP position obtained by forward kinematics solution of the joint angle data from the perspective of the theoretical model, which is not consistent with the actual position of TCP" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002609_dese.2019.00022-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002609_dese.2019.00022-Figure1-1.png", + "caption": "Fig. 1. Proposed quasi-omnidirectional robotic platform. a) Kinematic parameters (top view). b) Underactuated three-motor quasi-omnidirectional robot concept.", + "texts": [ + " The number of independent control variables is greater in number than the DOF in the working space. Usually walker robots control is devoted for cyclod gaiting control [13], resembling biomimetic gaits [18], gaits synthesis [5], synthesis for limbs reconfiguration [4], gait planning [10] and learning [11]. In this work\u2019s concept, an asymmetric hexapod deploying 64 978-1-7281-3021-7/19/$31.00 \u00a92019 IEEE DOI 10.1109/DeSE.2019.00022 Authorized licensed use limited to: Carleton University. Downloaded on June 30,2020 at 13:29:47 UTC from IEEE Xplore. Restrictions apply. Klann limbs is proposed (Figure 1a). The Klann linkage is a planar mechanism with 7 passive joints, using one rotary actuator, but in this work the mechanism is also be controlled in yaw. One of our work\u2019s main contributions is the mechanical design that reduced a traditional hyoer-redundantn eighteenservo model into an asymmetric three-motor underactuated version of hexapod, still preserving omnidirectional mobility (Figure 11b). One servomotor for limbs bidirectional synchronized steering, a second motor for driving the right-sided triplet of legs, and a third motor to drive the left-sided triplet", + " (21) Likewise, the inverse algebraic solution is expressed by \u03a6\u0307 = \u23a1 \u23a3 l1c0cs + l3c2cs + l4c3cs + l7c6cs \u2212l1s6ss \u2212l3s2ss \u2212l4s3ss \u2212l7s6ss 0 l1c0 l3c2 l4c3 l7c6 \u2212l1c0ss \u2212 l3c2ss \u2212 l4c3ss \u2212 l7c6ss \u2212l1s0cs \u2212l3s2cs \u2212l4s3cs \u2212l7s6cs \u23a4 \u23a6\u22121 \u00b7 p\u0307, (22) where the pseudoinverse matrix J+ = (J \u00b7 J )\u22121 \u00b7 J is an invertible, non singular and non stationary matrix. Moreover, the robots lateral speeds are defined by vr = v5 + v1 2 = \u2016q\u03075 + q\u03071\u2016 2 , o bien vl = \u2016q\u03074\u2016 (23) and vl = v2 + v6 2 = \u2016q\u03072 + q\u03076\u2016 2 , or vl = \u2016q\u03073\u2016 (24) From Figure 1a, the robot\u2019s angular velocity is described by \u03c9 = 2b(vr \u2212 vl) a2t + b2t (25) Considering that at and bt are non constant as legs yaw \u03c6s changes, where at = ||q5 \u2212 q1|| and bt = |zr \u2212 zl|. Moreover, zr = E cos(\u03c6s) and zl = E cos(\u03c6s), where E is the instantaneous length from the Klanns rotary point to the endeffectors contact point. When the angle \u03c6s = 0, then bt keeps constant and such an angle is aligned to robots longitudinal axis, and producing as a result the following specific case: \u03c9 = 2|zr \u2212 zl|( (\u2016q\u03075\u2016+\u2016q\u03071\u2016)\u2212(\u2016q\u03072+\u2016q\u03076\u2016) 2 ) (\u2016q5\u2016 \u2212 \u2016q1\u2016)2 + |zr \u2212 zl|2 (26) Therefore, the control vector u\u0307 is algebraically deduced with time-variant control matrix K,[ v \u03c9 ] = ( vr 2 + vl 2 2b a2 t+b2t (vr \u2212 vl) ) = ( 1 2 1 2 2b a2 t+b2t \u2212 2b a2 t+b2t ) \ufe38 \ufe37\ufe37 \ufe38 K \u00b7 ( vr vl ) (27) In addition, the functional forms of vr and vl are established by the general expression \u03bbq(\u03c60, \u03c6s) , which is in terms of the robots yaw orientation, such that ( vr vl ) = \u03bbt \u00b7 \u239b \u239c\u239d \u2016q\u03071(\u03c6\u03070, \u03c6\u0307s)\u2016 " + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001311_chicc.2019.8865180-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001311_chicc.2019.8865180-Figure1-1.png", + "caption": "Fig. 1: Basic strucuture of the quadrotor helicopter", + "texts": [ + " We define aij = 1 as information flows from quadrotor i to quadrotor j and 0 otherwise. If aij = aji = 1, then the quadrotor i and j can get information from each other simultaneously. Let G\u0304 be the graph that contains both graph G and one additional node for the leader quadrotor. We assume graph G is connected. G\u0304 contains a directed spinning tree rooted at the leader. More concretely, every follower quadrotor is connected to the undirected network of followers, while at least one follower quadrotor is connected to the leader. The basic structure of a quadrotor is shown in Fig. 1. To describe the attitude and the position, two coordinate sys- tems are established: body-fixed frame and inertial frame. The origin of body-fixed frame is located on center of mass and move with the aircraft. The inertial frame is fixed on the earth. Two coordinate systems coincide before taking off. We use Euler angles to describe the rotations. The rotations around x, y, and z-axis are denoted by \u03c6, \u03b8, and \u03c8, respectively. The quadrotor is actuated by rotors, which are located at a distance R from the center" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001999_14484846.2020.1714352-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001999_14484846.2020.1714352-Figure4-1.png", + "caption": "Figure 4. A vector loop of the ith active leg of the 4RSS+PS PM.", + "texts": [ + " Point C of the moving platformmoves along the z axis, so its position can be expressed by h (Figure 2). In addition, orientation of the moving platform was defined by three Euler angles \u03c6x, \u03c6y and \u03c6z. Therefore, pose (position and orientation) of the moving platform with respect to the fixed frame O-xyz can be fully defined by pose vector \u03b7mp \u00bc \u00bd h \u03c6x \u03c6y \u03c6z T (4) In inverse position kinematics of the manipulator, pose vector of the moving platform \u03b7mp is known, and rotation angles \u03b8i of actuators are to be calculated. Regarding Figure 4, a loop closure equation can be written for the ith RSS leg as follows hk \u00fe bri eik ai \u00bc lif i (5) where ai is given in terms of \u03b8i by ai \u00bc OiAi ! \u00bc \u00bd ai cos \u03b8i ai sin \u03b8i 0 T (6) Let rmp;i be representative of unit vector ri in the local frame C\u2013xmpympzmp, so we can write ri \u00bc Rmprmp;i (7) where r1;mp \u00bc \u00bd cos \u03b1 sin \u03b1 0 T (8a) r2;mp \u00bc \u00bd cos \u03b1 sin \u03b1 0 T (8b) and from the structure of the manipulator, we can find that rmp;3 \u00bc rmp;1 (8c) rmp;4 \u00bc rmp;2 (8d) Substituting ai and ri from Equations (6) and (7) into Equation (5), and squaring both sides of the resultant equation yields D1i sin \u03b8i \u00fe D2i cos \u03b8i \u00fe D3i \u00bc 0 (9) where coefficients D1i, D2i and D3i depend on the known kinematic parameters of the manipulator, as follows D1i \u00bc 2baiRmp;2 \u00f01 3\u00dermp;i (10a) D2i \u00bc 2baiRmp;1 \u00f01 3\u00dermp;i (10b) D3i \u00bc 2b\u00f0h ei\u00deRmp;3 \u00f01 3\u00dermp;i \u00fe b2 \u00fe a2i l2i \u00fe e2i 2eih\u00fe h2 (10c) In this paper, Ai \u00f01 n\u00de (for i = 1, 2, \u2026, m) denotes the ith row of a m \u00d7 n matrix A" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure60.12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure60.12-1.png", + "caption": "Fig. 60.12 Abrasive wear mechanism of GF Nylon 6 [8]", + "texts": [ + " values converge quickly. The best solution obtained is the objective function value of 0.5030 for glass-filled content of 30 wt%, sliding distance 500 m and load of 20 N for lower COF and SWR as shown in Figs. 60.9, 60.10 and 60.11. 60.3.4 Confirmation Tests for CSA With the identified optimal process parameters of glass-filled content of 30 wt%, sliding distance of 500 m and load of 20 N, a confirmation experiment was conducted as shown in Table 60.5. Fig. 60.10 Variation of sliding distance with iterations Figure 60.12 shows a schematic illustration of abrasive wear mechanism. Figure 60.13a shows typical micrograph of unabraded counter-face of 320 grit size SiC abrasive paper. In case of low sliding distance, repetition of the pin over the wear track is less so clogging of abrasive wear debris is observed low as shown in Fig. 60.14a. However, in the case of 500-m sliding distance repetition of the pin over the wear track is more resulting in less effect of abrasiveness as shown in Figs. 60.13b and 60.14b. Consequently, decrease COF and SWR" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003380_012027-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003380_012027-Figure1-1.png", + "caption": "Figure 1. The basic scheme of boiler KTU750: 1 \u2013 heat exchanger; 2 \u2013 cleanout of the heat exchanger; 3 \u2013 furnace door; 4 \u2013 burner; 5 \u2013 grate; 6 \u2013 fuel supply mechanism.", + "texts": [ + "1088/1757-899X/866/1/012027 September 2017 to April 2018, 742000 kW were spent on heating the administrative and amenity building, which is 2 619 260 rubles. In order to reduce heating costs, it was proposed to reconstruct the heating system of the administrative and amenity building by replacing the heat of an electric boiler with the heat energy of a wood waste boiler. For the normal operation of the boiler and the prospect for expansion of the company, as well as for the possible prospects of introduction of heat generation into the market, it was decided to purchase a boiler unit KTU-750 manufactured by \u201cTeploresurs\u201d LLC. Figure 1 shows a schematic diagram of KTU-750 boiler. The design feature of the furnace is an arch vault allowing highly moist fuel up to 55% to be burnt and it is made with one furnace front, which provides maintenance of the grate and loading of lump fuel (cuttings, firewood, etc.). The design of the boiler KTU-750 provides a window for the mechanized supply of bulk fuel (sawdust, wood chips, bark, peat, etc.). combustion. The heat exchanger with smoke tubes. The heat carrier (water) is pumped into the heat exchanger, flue gases wash the inner walls of the pipes where the heat exchange takes place" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000925_b978-0-12-816634-5.00003-0-Figure3.8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000925_b978-0-12-816634-5.00003-0-Figure3.8-1.png", + "caption": "Figure 3.8 Injection mold core using DMLS technology. (Left) CAD design showing internal cooling channels. (Right) DMLS-built actual core using 300 Maraging steel [29]. Courtesy: EOS North America, Adam J. Penna.", + "texts": [ + " Building the integrated mount and manifold with internal passageways in a single operation eliminates multiple part fabrication and results in significant cost savings. The good surface finish of the part eliminated finish machining needs on all surfaces, except seal surfaces and threading of screw holes. Generally the PBF technique gives a better surface finish than the DED approach, 63Comparison of various additive manufacturing technologies however for demanding applications (such as in aerospace) finish machining and/or other surface finishing operations are still required [28]. Fig. 3.8 shows an injection molding core built using DMLS technology [29]. This tool represents one core in a 16-core production tool for injection molding plastic parts. The left side shows a CAD model of the tool with internal cooling channels and the right side shows the actual tool built using 300 Maraging steel. The ability of PBF technologies to produce three-dimensional cooling passages following the part contours allows such tool building. One benefit of this is the placement of cooling channels closer to the part surface, leading to faster and uniform plastic cooling, which results in a shorter cycle time and better part quality" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001962_j.mechatronics.2020.102323-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001962_j.mechatronics.2020.102323-Figure3-1.png", + "caption": "Fig. 3. Diver robot.", + "texts": [ + " Step 4: Extract control U ( t ) from x \u2217 ( t ) using (21) Step 5: Integrate the dynamics (7) with extracted control U ( t ) and initial value x \u2217 (0) to obtain the integrated path ?\u0303? ( \ud835\udc61 ) , which is the planned motion. The same procedures apply for the motion planning with states contraints, where the AGHF (10) is replaced by the constrained AGHF (18) . . Implementation and simulation results .1. Diver robot We consider a planar diver robot with three links with revolute oints, as illustrated in Fig. 3 . The middle link, which we consider to e the base link, can be thought of as corresponding to a human torso. he position and orientation of base link in an inertial frame are denoted y [ x 0 , y 0 ] \u22a4 and \ud835\udf030 . Link 1 and link 2 are connected to opposite ends of he base link, and can be thought of the arms and legs. The relative anle between link 1, link 2 and the base link are q 1 and q 2 . The dynamics ollows (6) , with \ud835\udc5e = [ \ud835\udf030 , \ud835\udc5e 1 , \ud835\udc5e 2 ] \u22a4 and joint torques as input \ud835\udc48 = [ \ud835\udc62 1 , \ud835\udc62 2 ] \u22a4. The system parameters are chosen to be proportional to human\u2019s orso, arms and legs, and are displayed in the table below" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003513_j.promfg.2020.08.087-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003513_j.promfg.2020.08.087-Figure5-1.png", + "caption": "Fig. 5. Finite element model of hot stretch bending of I titanium profile.", + "texts": [ + "5Al-2Zr-1Mo-1V during hot stretch bending process and the material properties such as Young\u2019s modulus (85680MPa), Poisson\u2019s ratio (0.32), specific heat (920J/(Kg* \u2103 )) and conductivity(11.8 W/(m*\u2103)) were cited from the reference data [5]. Chen Zhang et al. / Procedia Manufacturing 50 (2020) 483\u2013487 485 C. Zhang et al. / Procedia Manufacturing 00 (2019) 000\u2013000 3 The coupled thermo-mechanical FE model of HSB was created in ABAQUS software, only semi-model was simulated because of the symmetry of model and the meshed model is presented in Fig.5. We adopted the dynamic, tempdisplacement, explicit procedure type in stretch bending analysis step while used the visco type in creep forming. The die and the simplified clamp were set as discrete rigid element (R3D4). And the profile was set as 3D stress element (C3D8R). And the tie constraint was applied between extrusion and the simplified clamp. The different uniform temperature fields were predefined to the extrusion, respectively. The clamp trajectory during stretch bending analysis step consisted of two steps , including pre-stretch and stretch bending" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002052_b978-0-12-812750-6.00001-9-Figure1.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002052_b978-0-12-812750-6.00001-9-Figure1.4-1.png", + "caption": "FIGURE 1.4 Load paths of vehicle frontal structure.", + "texts": [ + " Research (Brantman, 1991) has shown that for optimal occupant safety in a collision at 48 km/h impact velocity, the first phase lasts between 10 and 30 ms, the second phase lasts 35 ms, and the last phase fills up the remaining time to a total of maximal 90 ms. The design process of the crashworthiness structure seeks to control the paths of load transformation and to optimize the energy-absorbing process and acceleration pulse. For example, there are several load paths in the frontal structure of a modern Sedan, as shown in Fig. 1.4: Path 1: Accessories\u2014bumper\u2014crash boxes\u2014longitudinal beams Path 2: Upper rails\u2014A pillar Path 3: Subframes\u2014sill beams The components in three paths are deformable and can absorb the impact energy. Notably, the first path absorbs more than 50% of the total crash energy in most frontal crashes (Gris\u030ckevicius and Z\u030ciliukas, 2003). Fig. 1.5 shows a rough estimation of energy absorption distributed on the different components of the frontal structure during a crash at 56 km/h against a rigid barrier. The features and functions of these components are as follows: Bumper: the bumpers are usually reinforcement bars made of steel, aluminum, plastic, or composite material and can absorb crash energy to a certain extent" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001597_j.cad.2019.102798-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001597_j.cad.2019.102798-Figure3-1.png", + "caption": "Figure 3: Bounding Box Test An idealized case of a single breakpoint placement for a quarter circle scenario is shown in A. In B, dashed bounding box are drawn.", + "texts": [ + " Bounding Box Breakpoints225 Each segment Si can be further split using a recursive bounding box approach. Assuming that Si consist of t number of points, {p1 . . . pt}, it is recursively split Si into two subsets at index k, where k \u2208 2 . . . t\u2212 1, if it fulfills two conditions: pk has the largest perpendicular distance to the line formed by p1 and pt, and the ratio of the perpendicular distance to the line exceeds 0.15. The230 rationale for using a criterion of 0.15 is based on the idealized case of a quarter circle corner, as illustrated in Figure 3A. An ideal placement would be a single breakpoint placed at the middle of the quarter circle, regardless of the length 10 Jo ur al Pr epr oo f from either the start or end to the quarter circle. In this ideal case scenario, a boundary box ratio of 0.15 is used to limit the unnecessary placement of new235 bounding box breakpoint as depicted in Figure 3B. Finally, the bounding box breakpoints, are added to the index set b and the updated set b is used as the index set for the fixed boundary points. Figure 4 shows the generation of the three types of breakpoints for the marine blade As mentioned earlier, the fixed boundary point set will always exist in every blocking solution. In addition, it is necessary to define the candidate point 11 Jo ur re set, which comprises of the boundary and interior candidate points, from which the EA can draw from to derive a possible blocking solution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000925_b978-0-12-816634-5.00003-0-Figure3.12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000925_b978-0-12-816634-5.00003-0-Figure3.12-1.png", + "caption": "Figure 3.12 (Left) Damaged tool; (middle) scanned image of the damaged tool; (right) rebuilt tool [33]. Courtesy: EOS North America, Adam J. Penna.", + "texts": [ + "2 m Ti-6Al-4V wing spar, which was deposited in a flexible enclosure using plasma arc welding with a seven-axis robotic system [32]. The part features straight and curved features, all printed perpendicular to the substrate. Two parts were built simultaneously by alternating deposition on either side of a sacrificial substrate in order to balance and minimize distortion stresses on the substrate plate. The deposition rate was 0.8 kg/h with a buy-to-fly ratio of 1.2. Repair and remanufacturing of worn out and damaged components is an important application area for AM. Fig. 3.12 shows a tool insert with one of the fingers on the top broken during service [33]. Once the tool was scanned and a CAD geometry regenerated, the tool was rebuilt using DMLS technology. Instead of replacing the damaged tool with a new tool, repair of the damaged tool can make significant savings. One of the best application areas suited for DED techniques is repair and remanufacturing. Due to their ability to add metal on selected locations on 3D surfaces, these technologies can be used to rebuild lost material on various components [34 36]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000790_iraniancee.2019.8786525-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000790_iraniancee.2019.8786525-Figure1-1.png", + "caption": "Fig. 1. The wind turbine components", + "texts": [ + " Fluctuations may cause fatigue and unwanted loading in mechanical components and actuators [5, 6]. Pitch mechanism should be simple to maintain, relatively compact, lightweight, and reliable [7]. In this paper after explaining the structure and components of a typical wind turbine with permanent magnet synchronous generator, an idea to overcome the pitch signal fluctuations is represented. This method is applied to the gain scheduled PI pitch controller of an experimental simulator for 100 kW wind turbine. II. WIND TURBINE COMPONENTS As it is shown in fig. 1, there are four main parts in a wind turbine: the base, tower, nacelle and blades. The blades capture the mechanical power from the wind [1, 7]. There are a generator and sometimes a gearbox in the nacelle. The gearbox of a typical wind turbine should be robust enough to manipulate the frequent variations in the mechanical torque originated by the wind speed variations. Generators can be either variable or fixed speed. Variable speed generators produce electricity at a varying frequency, which must be changed to 50 or 60 Hz before it is fed in to the electrical grid [8]" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001917_iceerp49088.2019.8956980-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001917_iceerp49088.2019.8956980-Figure4-1.png", + "caption": "Fig. 4. Magnetic flux density distribution (I = 3.5 Arms, Nr = 100,000 rpm).", + "texts": [ + " Iron loss characteristics of the magnetic composite material are indicated in Fig. 2(b). We measured these characteristics using the inductor cancelation method. Moreover, the winding conditions are given in Fig. 3. We conducted the subsequent comparisons according to this dimension. (1) Comparison of coreless motor and IWSM characteristics with tc = 0.32 and wc = 1.0 mm and; (2) Loss comparison when changing to tc = 0.32, 0.30, 0.28, 0.26 mm and wc = 1.0, 0.9, 0.8, 0.7 mm in IWSM IV. NUMERICAL RESULTS AND DISCUSSION Fig. 4 shows the magnetic flux density distribution. By arranging the magnetic composite material around the winding, we observed that the magnetic flux passed through the magnetic composite and the magnetic flux linked to the winding was reduced. In addition, concentration of the magnetic flux on the magnetic composite material resulted to an increase in the gap magnetic flux density. Fig. 5 shows the current density distribution. By arranging the magnetic composite material, we observed that the magnetic flux linking the windings was reduced, together with the deviation of the current density distribution" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001026_iccsdet.2018.8821181-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001026_iccsdet.2018.8821181-Figure1-1.png", + "caption": "Fig. 1. Cantilever Beam Geometry with highlighted ground Plane.", + "texts": [ + "42, g0 is initial gap between the beam and ground, 0 is Permittivity of free space, L is the length of the beam and W is the width of the beam. The expression [6] for B is given by Equation (2) where \u00ca is the young\u2019s modulus of the cantilever beam and H is the beam thickness. The Young's modulus of each layer should be taken into account when deriving the equivalent Young\u2019s Modulus Ee of the composite beam [7]. where En, tn are the Young's modulus and thickness of the each layer respectively and is the total thickness of the beam. The Geometry of the Cantilever beam is shown in Fig. 1 and Cantilever Beam dimensions are shown in Table I. COMSOL Simulation is performed for different Bimorph Cantilever Beam Materials. Gold, Copper, Titanium and Molybdenum materials are used as one of the Bimorph Layer and Aluminum as the common bimorph Material. Thermal expansion Coefficient for Aluminum is greater than Gold,copper, Titanium and Molybdenum material. In Electrostatic Actuation, electrostatic force bend the cantilever beam and thereby reduce the gap between the beam and ground plane" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000726_978-3-030-21013-7_15-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000726_978-3-030-21013-7_15-Figure3-1.png", + "caption": "Fig. 3 An illustrative rigid tool-flexible workpiece milling", + "texts": [ + " For upmilling, \u03b8s = 0 and \u03b8e = cos\u22121(1 \u2212 2\u03c1)while for down-milling, \u03b8s = cos\u22121(2\u03c1 \u2212 1) and \u03b8e = \u03c0 where \u03c1 = B/D is the radial immersion. A illustration of this model showing the geometric and modal parameters is shown in Fig. 1. On inserting the above model matrices in the constructed monodromy matrix and substituting the numerical values given in Table 1, the stability diagrams given in Fig. 2 are computed. The results agree with the known results in [4, 23]. This milling case is illustrated in Fig. 3 showing the primary motion (spindle rotation ) and the secondary motion (feed v) of the tool. The thin-walled workpiece is much more flexible than the tool and it is thus considered compliant while the tool is considered rigid. Since the workpiece is a continuum, a large DOF second order delayed model is needed to capture the regenerative dynamics. The regenerative dynamics of the thinwalled workpiece is, therefore, subjected to Finite Element (FE) modelling to give a dF -dimensional model for every location of the tool along the feed direction" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003059_b978-0-12-821350-6.00012-3-Figure11-1.png", + "caption": "Fig. 11 CAD drawing of the missing rib design.", + "texts": [ + " 10) can be classified as a hybrid of the chiral and re-entrant designs. This structure is formed by taking the base hexagonal honeycomb re-entrant design and removing two opposite ribs out of the four. This is done because unit cells with four ribs are much stiffer than those with two ribs, and, as such, can display auxetic behavior at much lower loads. The missing rib model has broadly two types of auxetic geometries, namely the Lozenge grid and the square grid. For our testing, we used the Lozenge grid design as shown in the CAD in Fig. 11. The in-plane Poisson\u2019s ratio for this geometry was found to be 0.43 (Kolken and Zadpoor, 2017). The detailed kinematic model of the missing rib configuration was presented by Zhai (Zhai et al., 2018). The auxetic effect of such a design is obtained by the concurrent out-folding of re-entrant cells and the missing ribs\u2019 rotation mechanism. The double arrowhead design is another variant of the re-entrant honeycomb, which was first founded through the numerical topology optimization method. Based on the actual configuration of the arrowhead, any extension will cause the triangles to expand in the transverse direction while compressions will cause them to collapse" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001617_j.ifacol.2019.11.189-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001617_j.ifacol.2019.11.189-Figure1-1.png", + "caption": "Fig. 1. Frames representation.", + "texts": [ + " This principle is known as Active Disturbance Rejection Control (ADRC). The numerical simulations have shown satisfactory results. The document is organized as follows: in Section 2, a simple control-oriented model is presented. Section 3 introduces the reference model-based control strategy. Section 4 shows the design of the proposed nonlinear booster and observer. Simulation 21st IFAC Symposium o Automatic Control in Aerospace August 27-30, 2019. Cranfield, UK Copyright \u00a9 2019 IFAC 152 \ud835\udf13\ud835\udf13 \ud835\udf03\ud835\udf03 \ud835\udf11\ud835\udf11 \ud835\udc4b\ud835\udc4b\ud835\udc35\ud835\udc35 \ud835\udc4d\ud835\udc4d\ud835\udc35\ud835\udc35 \ud835\udc4c\ud835\udc4c\ud835\udc35\ud835\udc35\ud835\udc42\ud835\udc42\ud835\udc35\ud835\udc35 \ud835\udc4b\ud835\udc4b\ud835\udc38\ud835\udc38 \ud835\udc4c\ud835\udc4c\ud835\udc38\ud835\udc38 \ud835\udc4d\ud835\udc4d\ud835\udc38\ud835\udc38 \ud835\udc42\ud835\udc42\ud835\udc38\ud835\udc38 Fig. 1. Frames representation. results are illustrated in Section 5. Conclusions are given in the last section. 2. CONTROL-ORIENTED MODEL Quadrotors are nonlinear, highly coupled and complex systems. However, simplified models are usually used in the literature that are considered as control-oriented models to obtain simple control laws. Moreover, through these simplified models, we investigate the effectiveness of the designed control strategies if they can handle the disturbances and the uncertainties. Hereafter, in low-speed flight conditions, the used controloriented model neglects some effects such as: blade-flapping moments, hub forces, gyroscopic moments, etc. These effects have a minor impact on the vehicle and will be gathered in one disturbance vector \u2206 that includes all the neglected and unmodeled dynamics, external disturbances, etc. The vehicle operates in two coordinate frames: the Earthfixed frame RE(OE ,XE ,YE ,ZE) and the Body-fixed frame RB(OB,XB,YB,ZB) (as shown in Figure 1). RB is attached to the vehicle and constrained to move with it. We then get x\u0308 = u1 c\u03c8 s\u03b8 c\u03d5+s\u03c8 s\u03d5 m +dx y\u0308 = u1 s\u03c8 s\u03b8 c\u03d5\u2212c\u03c8 s\u03d5 m +dy z\u0308 =\u2212g+u1 c\u03b8 c\u03d5 m +dz \u03d5\u0308 = (Iy \u2212 Iz) Ix \u03b8\u0307 \u03c8\u0307 + u2 Ix +d\u03d5 \u03b8\u0308 = (Iz \u2212 Ix) Iy \u03d5\u0307\u03c8\u0307 + u3 Iy +d\u03b8 \u03c8\u0308 = (Ix \u2212 Iy) Iz \u03d5\u0307\u03b8\u0307 + u4 Iz +d\u03c8 (1) where m is the mass and I = diag(Ix, Iy, Iz) is the inertia matrix about the center of gravity. g denotes the gravitation coefficient. Let \u03b7 = (\u03d5,\u03b8 ,\u03c8)T \u2208 R3 be the orientation (Roll, Pitch, Yaw) of the quadrotor and let \u03c7=(x,y,z)T \u2208 R3 denoting its absolute position with respect to RE with \u03d5 = \u03c0 2 +k\u03c0,\u03b8 = \u03c0 2 +k\u03c0,k \u2208Z", + " Moreover, through these simplified models, we investigate the effectiveness of the designed control strategies if they can handle the disturbances and the uncertainties. Hereafter, in low-speed flight conditions, the used controloriented model neglects some effects such as: blade-flapping moments, hub forces, gyroscopic moments, etc. These effects have a minor impact on the vehicle and will be gathered in one disturbance vector \u2206 that includes all the neglected and unmodeled dynamics, external disturbances, etc. The vehicle operates in two coordinate frames: the Earthfixed frame RE(OE ,XE ,YE ,ZE) and the Body-fixed frame RB(OB,XB,YB,ZB) (as shown in Figure 1). RB is attached to the vehicle and constrained to move with it. We then get x\u0308 = u1 c\u03c8 s\u03b8 c\u03d5+s\u03c8 s\u03d5 m +dx y\u0308 = u1 s\u03c8 s\u03b8 c\u03d5\u2212c\u03c8 s\u03d5 m +dy z\u0308 =\u2212g+u1 c\u03b8 c\u03d5 m +dz \u03d5\u0308 = (Iy \u2212 Iz) Ix \u03b8\u0307 \u03c8\u0307 + u2 Ix +d\u03d5 \u03b8\u0308 = (Iz \u2212 Ix) Iy \u03d5\u0307\u03c8\u0307 + u3 Iy +d\u03b8 \u03c8\u0308 = (Ix \u2212 Iy) Iz \u03d5\u0307\u03b8\u0307 + u4 Iz +d\u03c8 (1) where m is the mass and I = diag(Ix, Iy, Iz) is the inertia matrix about the center of gravity. g denotes the gravitation coefficient. Let \u03b7 = (\u03d5,\u03b8 ,\u03c8)T \u2208 R3 be the orientation (Roll, Pitch, Yaw) of the quadrotor and let \u03c7=(x,y,z)T \u2208 R3 denoting its absolute position with respect to RE with \u03d5 = \u03c0 2 +k\u03c0,\u03b8 = \u03c0 2 +k\u03c0,k \u2208Z" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003657_j.matpr.2020.08.023-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003657_j.matpr.2020.08.023-Figure2-1.png", + "caption": "Fig. 2. The topology of 1-DOF manipulator-link.", + "texts": [ + " The MATLAB code dynamically updates the design variables based on the selection volume fraction as 0.5. The code can capture performance values such as deflection and Von-Mises stress. For smoothing the curves or boundaries of the topology, the MATLAB code is also incorporated with mesh independence, grayscale removal filters, and sensitive analysis [16]. The manipulator-link is made up of Aluminium of density 2700 kg/m3, and 1.5 kg payload is acting over it. The obtained topology for the 1-DOF link based on stated problem is shown in Fig. 2. The performance values of the solid link (volume fraction 1.0) and optimum topology at volume fraction 0.5 are shown in Table 1. When volume fraction is reduced to 0.5, the deflection and VonMises stress are increased by 59.5% and 42.4%, respectively. However, the performance values of the manipulator-link are within acceptable limits. The obtained topology from MATLAB result is imported to SOLIDWORKS software to convert topology from 2D to 3D at the desired thickness. Computer numerical control milling machine is used to manufacture the optimum link with the help of G codes developed in MASTERCAM software" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001361_978-3-030-30655-7_2-Figure2.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001361_978-3-030-30655-7_2-Figure2.2-1.png", + "caption": "Fig. 2.2 STA phase trajectory", + "texts": [ + " Figure or Figures FNN Fuzzy neural network FOU Footprint of uncertainty FPGA Field-programmable gate array FTC Fault tolerant control GDM Gradient descent method HGO High-gain observer HIL Hardware-in-loop HOSM High-order sliding mode ICE Internal combustion engine IGBT Insulated gate bipolar transistor IT2 Interval type-2 IT2FNN Interval type-2 fuzzy neural network KF Kalman filter kVAr Kilovolt-amperes reactive, thousand Volt-ampere reactive, a unit of reactive power LMF Lower membership function LMI Linear matrix inequality xviii Notations and Acronyms LPV Linear parameter varying LTI Linear time-invariant MCFC Molten carbonate fuel cell PAFC Phosphoric acid fuel cell PEM Proton exchange membrane PEMFC Proton exchange membrane fuel cell PR Proportional plus resonant PWM Pulse-width modulation THD Total harmonic distortion T-S Takagi-Sugeno SMC Sliding mode control SMO Sliding mode observer SOFC Solid oxide fuel cell SOSM Second-order sliding mode SOSML The addition of linear term to the nonlinear SOSM term SPD Symmetrical positive definite SRF Synchronous reference frame ST Super-twisting STA Super-twisting algorithm UIO Unknown input observers UKF Unscented Kalman filter VSC Variable structure control Notations and Acronyms xix List of Figures Fig. 1.1 PEMFC power system . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2 Fig. 1.2 Observer as the heart of control systems [2] . . . . . . . . . . . . . . . . 3 Fig. 1.3 Basic configuration of observer-based FDI [14] . . . . . . . . . . . . . 5 Fig. 1.4 Main contents of this publication . . . . . . . . . . . . . . . . . . . . . . . . 7 Fig. 2.1 SOSM trajectory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 Fig. 2.2 STA phase trajectory . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 21 Fig. 2.3 Twisting algorithm phase trajectory . . . . . . . . . . . . . . . . . . . . . . 22 Fig. 2.4 Sub-optimal algorithm phase trajectory. . . . . . . . . . . . . . . . . . . . 23 Fig. 3.1 Frequency characteristics of STA . . . . . . . . . . . . . . . . . . . . . . . . 36 Fig. 3.2 Frequency characteristics of SOSML . . . . . . . . . . . . . . . . . . . . . 38 Fig. 3.3 Estimates of state and disturbance . ", + " Given that the expression for the sliding manifold is known, it is possible to design the constant parameters of the controllers [31]. Super-Twisting Control Algorithm The super-twisting algorithm (STA) is a unique absolutely continuous sliding mode algorithm ensuring all the main properties of first order sliding mode control for system with Lipschitz continuous matched uncertainties/disturbances with bounded gradients [18]. The STA was developed to control systems with relative degree one in order to avoid chattering in variable structure control (VSC). The trajectories on the second sliding manifold are shown in Fig. 2.2. Consider the system (2.23), the control algorithm is defined as follows [25] where \u03b1, \u03bb are positive constants and \u03c1 \u2208 (0, 1). The sufficient conditions for the finite time convergence to the sliding manifold are \u03b1 > C Km , \u03bb2 \u2265 4C K 2 m KM(\u03b1 + C) Km(\u03b1 \u2212 C) . (2.28) The STA does not need the evaluation of the sign of the time derivative of the sliding variable. For the choice \u03c1 = 1, the origin is an exponentially stable equilibrium point. The choice \u03c1 = 0.5 assures that the maximum real second order sliding is achieved" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure8.1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure8.1-1.png", + "caption": "Fig. 8.1 Through-thickness crack of length 2c in a hollow cylinder subject to internal pressure. How does one experimentally test such a case?", + "texts": [ + " Like all pressure vessels, the pipes must satisfy the \u2018Leak before break\u2019 criterion to ensure that the reactor can be safely shut down in the unlikely event that a hydride-induced crack grows through the wall thickness, resulting in leakage. Nuclear power reactors are equipped with moisture detectors to close this safety loop. BISS was contracted to design, develop, manufacture, validate and supply a test rig to evaluate the residual (burst) strength of controlled size sections of (irradiated) piping from pressurized water reactors. A detailed description of this project is given in Avinash et al. (2006). Here we present the technical details of interest to the experimental research community. Figure 8.1 is a schematic of a hollow cylinder with an axially-orientated throughthickness crack. \u2018Handbook\u2019 solutions are available for the stress intensity factor, K, for a crack in the presence of internal pressure. Also available is the equation that describes the open area formed by the crack with such a pressurized component or specimen geometry. Assuming the crack to take the shape of an ellipse under load, one can derive the equation for crack opening displacement under internal pressure and proceed to describe the compliance function. Thus, in principle, a specimen simulating the configuration shown in Fig. 8.1 can serve the purpose of fracture testing, just like the ASTM standard compact-tension and single edge-notched specimens used by industry and the research community for more general fracture mechanics testing. However, unlike the solutions and equations for ASTM standard specimens, the solutions and equation for a pressure vessel may never have been experimentally verified, owing to the difficulties mentioned in the next paragraph. In contrast to standard laboratory coupons, testing a hollow cylinder for fracture presents several challenges: \u2022 The through-crack needs to be sealed in a manner that satisfies contradictory requirements: (i) it must sustain internal pressure without leakage and (ii) at the same time permit the crack to open and close, firstly to permit fatigue pre-cracking for a valid J1c test; and secondly to permit stable crack growth and automatic crack size (and crack growth increment) measurements during the fracture test" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000546_978-3-030-20131-9_278-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000546_978-3-030-20131-9_278-Figure3-1.png", + "caption": "Fig. 3. Kinematic schemes: (a) mechanism 1a,2,3,6 for the rotation of link 6 around the axis defined by points P and Q, (b) mechanism 1c,4,5,6 for the rotation of link 6 around the axis defined by points R and S, and (c) mechanism 6a,9,10,8 for the rotation of link 8 in relation to link 6 around the axis defined by the points O1 and O2.", + "texts": [ + " In addition, the axes of the coordinate systems O2x2 1y2 1z2 1 and O2x2 2y2 2z2 2 are parallel with the axes of the coordinate systems O1x1 1y1 1z1 1 and O1x1 2y1 2z1 2. Due to the transparency of the Fig 2b, the coordinate system O2x2 1y2 1z2 1 is not shown. Finally, after rotation around the z2 2 axis for the angle 8 6, the coordinate system O2x2 2y2 2z2 2 switches to the position O2x2 3y2 3z2 3. Following, based on known displacements of the input links 2, 4 and 9, by applying direct kinematics, rotation angles and angular velocities of the output links 6 and 8 are determined. Figure 3a shows the mechanism 1a,2,3,6 for the rotation of link 6 around the axis defined by points P and Q. The rotation angle of link 6 is determined according to: ( ) ( )( ) ( ) ( ) 2 22 1 0 2 1 0 2 6,1/2 2arctan OO VA s OO VA s A AB OV A AB OV \u03d5 \u2212 + \u00b1 \u2212 + \u2212 + + = \u2212 + (1) where: ( ) ( )( )222 2 1 0 2 1 12 BF AB OV OO VA s O F A O F \u2212 + \u2212 \u2212 + \u2212 = (2) The angular velocity of link 6 is determined according to: ( ) 2 6 1 6 6 3cos sin cot s O F \u03d5 \u03d5 \u03d5 \u03d5 = \u2212 (3) where: 1 6 3 cos arccos O F AB OV BF \u03d5\u03d5 \u2212 \u2212 = (4) Figure 3b shows the mechanism 1c,4,5,6 for the rotation of the link 6 around the axis defined by points R and S. The rotation angle of link 6 is determined according to: 2 2 2 6,1/ 2 2arctan\u03c8 \u2212 \u00b1 + \u2212= \u2212 F E F G G E (5) where: ( ) [ ]( )TE = \u2212 \u2212 \u22120 ub 1 0a b I Q b b (6) ( ) [ ]( )TF = \u2212 \u22120 ub 1 0a b P b b (7) ( ) [ ]( ) ( ) ( ) ( ) ( ) ( ) ( ) T T T T 1 2 1 1 2 2 G = \u2212 \u2212 + \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 \u2212 0 ub 1 0 1 1 1 1 0 0 1 0 1 0 a b Q b b a b a b a b a b b b b b (8) where: 1 0 4 1 0 1 0 0 ; 0 ; 0 ; 0 OU CD OU CD O E UC s OO UC OO + + = = = = + 0 1 1a b a b (9) [ ] [ ] 2 2 6 2 6 0 0 0 ; ; cos 0 sin bx bx by bx bzbz by bx bz bx bx by by by bz by by bx bzbx bz by bz bz u u u u uu u u u u u u u u u u u u uu u u u u \u03d5 \u03d5 \u2212 = = \u2212 = = \u2212 \u2212 = \u2212 ub ubP Q (10) The angular velocity 6\u03c8 is determined according to: ( ) ( ) ( ) ( ) ( ) 6 T T 6 T 0 0 bx by bz u u u \u03d5 \u03c8 \u2212 \u2212 \u2212 \u00d7 \u2212 = \u2212 \u00d7 \u2212 0 0 a b a a b b b a b b b (11) where: ( )6 6 6 4 0 0 ; ; 1 cosV s \u03c8 \u03c8 = = \u2212 + = \u2212 ,ub 1 0 0a b R b b b (12) 2 6 6 6 6 6 6 2 6 6 6 6 6 6 6 2 6 6 6 6 6 6 cos sin sin sin cos sin sin sin cos bx bx by bz bx bz by bx by bz by by bz bx bx bz by by bz bz bz u V u u V u u u V u u u V u u V u u V u u u V u u u V u u V \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 \u03c8 + \u2212 + = + + \u2212 \u2212 + + ,ubR (13) Link 6 first rotates for the angle 6, and then for angle 6. By differentiating these two angles, angular velocities are obtained whereby the angular velocity 6\u03d5 is pointed in the direction of the x axis, and 6\u03c8 is pointed in the direction of the axis defined by the points R and S, i.e. by its unit vector ub \u2013 see Fig. 3b. The total angular velocity of link 6 ( md) is defined as the vector sum of the angular velocities: 6 6 6 60 0 bx by bz u u u \u03d5 \u03d5 \u03c8 \u03c8 = + = + md bi u (14) Figure 3c shows the mechanism 6a,9,10,8 for the rotation of link 8 relative to link 6 around the axis defined by points O1 and O2. The rotation angle of link 8 relative to link 6 is determined according to: ( ) ( )22 2 0 9 2 0 9 16 8,1/ 2 1 2 2 arctan M L s LO M L s A A LO \u03b8 \u2212 \u2212 \u00b1 + \u2212 \u2212 = + (15) where: ( )22 2 2 2 0 9 2 1 22 O L LM s O N NM A O N + \u2212 + \u2212 = (16) The angular velocity of link 8 relative to link 6 is determined according to: ( ) 6 9 8 6 6 2 8 8 10cos sin cot s O N \u03b8 \u03b8 \u03b8 \u03d5 = \u2212 (17) where: 6 2 8 2 10 cos arccos O N O L NM \u03b8\u03d5 \u2212 = (18) Based on the kinematic analysis of the waist mechanism, the kinematic model was formed and motion simulation for initial movements of flexion, extension, lateral flexion and rotation \u2013 see Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002173_icrera47325.2019.8996549-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002173_icrera47325.2019.8996549-Figure10-1.png", + "caption": "Fig. 10. Electron microscope micrograph of the established passive temperature threshold sensor. The thermal actuator is combined with a bistable beam and two latching mechanisms. The first one allows to activate the device at a well defined activation temperature and the second latch establishes an irreversible contact when an upper threshold temperature is reached.", + "texts": [ + " For the proof-of-concept, the actuator is combined with a bistable switch to establish a passive temperature threshold sensor. Such a sensor allows for the tracking of proper storage and transportation temperatures of cooling goods, e.g., refrigerator cargo or blood bottles. The bistable beam switch is described in detail in [28] and [29]. It exhibits a total length of lbb = 2600 \u03bcm, an effective apex height of ha = 33 \u03bcm, a beam width of t = 7 \u03bcm, and is manufactured with the same process as the actuator. Two latching mechanisms are attached to the actuator and the bistable beam (Fig. 10). At first, the temperature is decreased until a well defined activation temperature is reached, where the first latch is activated. When the temperature is increased afterwards, the actuator pushes against the bistable beam until an upper threshold temperature is reached. At this temperature, the bistable beam switches into its second stable position and an irreversible contact is established with the second latching mechanism. A deflection of at least dmin = 44 \u03bcm and a peak force of fmax = 100 \u03bcN is required to push the bistable beam into its second stable position, where the available temperature change is only in the order of several 10 K" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure73.7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure73.7-1.png", + "caption": "Fig. 73.7 Velocity profiles at bends of serpentine microchannels", + "texts": [ + " Sharma higher Reynolds numbers, mixing of fluid was more to carry out more amount of heat from the substrate. However, at any particular Reynolds number, the base temperature of the V-serpentine microchannel was more. This attributes that the cooling effect on the V-serpentine was less compared to the other two serpentine channels. The reason could be that due to the presence of sharp corners in the V-serpentine, the possibility of fluid stagnation was more which results in higher base temperatures. 73.4.3 Velocity Profile Figure 73.7 indicates the velocity profiles of all the three serpentine microchannel bends. From Fig. 73.7, it is clearly visible that the more amount of fluid was stagnated at the sharp corners of the V-serpentine followed by rectangular and least in the U-serpentine microchannel. Therefore, the contribution of the fluid in these regions in heat transfer was very low. As the stagnant fluid region was high in V-serpentine, this leads to higher base temperatures. However, the stagnation region of fluid in U-serpentine was very less and hence lowers substrate base temperature (Fig. 73.6). 73.4.4 Heat Transfer Coefficient The heat transfer coefficient of the fluid passing through the microchannels was calculated using Eq" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002029_iccas47443.2019.8971518-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002029_iccas47443.2019.8971518-Figure3-1.png", + "caption": "Fig. 3 Detailed structure of expandable drill bit.", + "texts": [], + "surrounding_texts": [ + "1276\nDrilling technology has been utilized in plenty of industrial fields such as collecting underground information and resource exploration. Recently, technology development is being carried out in order to utilize it in an extreme area and space environment. The world is concentrating on exploring and collecting new energy resources such as shale gas, coalbed methane and Rare Earth Elements (REE). REE is continuously growing in demand due to the development of electronic devices, however, there are difficulties in forming a market because explored of buried area and supply are limited [1]. The drilling technique is classified into shallow drilling which excavates shallow depths and deep drilling which excavates deep depths. Most of the new energy sources mentioned are found to be distributed in the shallow drilling range [2]. Rotary Steering System (RSS), which is currently being applied, has the advantages of reducing drilling period and cost compared with vertical drilling systems [3-4]. On the other hand, it has several limitations such as applicable topography and cause of environmental pollution because of its enormous size and soil removing process. In addition, existing drilling systems must be accompanied by ancillary equipment such as rigs and water circulation systems. Additionally, the deeper the depth of drilling, the more complicated the process, such as the connecting supplementary pipes, resulting in heavy manpower and cost. In recent years, research on directional drilling has been underway, but it is not much different from existing vertical drilling systems, and there are no cases applied to the actual fields. Honeybee robotic in the United States and Japan Aerospace Exploration Agency (JAXA) are developing directional drilling technology for space exploration, and one of the projects in Europe, named BADGER is developing directional drilling robot that combines\nearthworm\u2019s interlocking motion and 3D printing technology [5-7]. Depending on the working environment, it is difficult to input equipment or manpower for drilling. In order to solve this problem, a necessity for a technique capable of minimizing cost and manpower by exploring various locations within a single borehole through directional drilling is needed. In recent years, bio-inspired or bio-mimetic robots have been actively developed to create mechanisms that mimic the movements and habits of living things [8-9]. The directional drilling robot proposed in this paper is an embedded system which can be drilled using only small-scaled drilling robot without large equipment and aims to improve the drilling efficiency compared to the existing method by mimicking the behavior of the mole which is living under the ground.\nIn this paper, moles, one of the mammals which are freely moving through the ground, looking for a feed, and living in their own dens with excavation habit, are simulated. The European mole uses two large paws for excavation and removing debris at the same time. The naked mole-rat scratches the ground with their incisors and moves the debris backward by using the front and rear limbs. These different types of moles with different excavation methods also relevant to the ground type. A naked mole-rat can open their mouth widely for a broad excavation area and have lip structures to prevent soil penetration [10-11]. Additionally, it has limbs for collecting and sweeping the excavated soil. It is possible to move more freely under the tunnel due to forming a wider borehole than the size of the body. This is possible owing to the high degree of freedom of the neck movement. These various excavation habits and\n978-89-93215-18-2/19/$31.00 \u24d2ICROS\nAuthorized licensed use limited to: Macquarie University. Downloaded on June 22,2020 at 19:52:00 UTC from IEEE Xplore. Restrictions apply.", + "1277\n2019 19th International Conference on Control, Automation and Systems (ICCAS 2019) Oct. 15~18, 2019; ICC Jeju, Jeju, Korea\n1. INTRODUCTION\nDrilling technology has been utilized in plenty of industrial fields such as collecting underground information and resource exploration. Recently, technology development is being carried out in order to utilize it in an extreme area and space environment. The world is concentrating on exploring and collecting new energy resources such as shale gas, coalbed methane and Rare Earth Elements (REE). REE is continuously growing in demand due to the development of electronic devices, however, there are difficulties in forming a market because explored of buried area and supply are limited [1]. The drilling technique is classified into shallow drilling which excavates shallow depths and deep drilling which excavates deep depths. Most of the new energy sources mentioned are found to be distributed in the shallow drilling range [2]. Rotary Steering System (RSS), which is currently being applied, has the advantages of reducing drilling period and cost compared with vertical drilling systems [3-4]. On the other hand, it has several limitations such as applicable topography and cause of environmental pollution because of its enormous size and soil removing process. In addition, existing drilling systems must be accompanied by ancillary equipment such as rigs and water circulation systems. Additionally, the deeper the depth of drilling, the more complicated the process, such as the connecting supplementary pipes, resulting in heavy manpower and cost. In recent years, research on directional drilling has been underway, but it is not much different from existing vertical drilling systems, and there are no cases applied to the actual fields. Honeybee robotic in the United States and Japan Aerospace Exploration Agency (JAXA) are developing directional drilling technology for space exploration, and one of the projects in Europe, named BADGER is developing directional drilling robot that combines earthworm\u2019s interlocking motion and 3D printing technology [5-7]. Depending on the working environment, it is difficult to input equipment or manpower for drilling. In order to solve this problem, a necessity for a technique capable of minimizing cost and manpower by exploring various locations within a single borehole through directional drilling is needed. In recent years, bio-inspired or bio-mimetic robots have been actively developed to create mechanisms that mimic the movements and habits of living things [8-9]. The directional drilling robot proposed in this paper is an embedded system which can be drilled using only small-scaled drilling robot without large equipment and aims to improve the drilling efficiency compared to the existing method by mimicking the behavior of the mole which is living under the ground. 2. BIOLOGICAL STRUCTURE AND DIGGING HABIT OF THE MOLE In this paper, moles, one of the mammals which are freely moving through the ground, looking for a feed, and living in their own dens with excavation habit, are simulated. The European mole uses two large paws for excavation and removing debris at the same time. The naked mole-rat scratches the ground with their incisors and moves the debris backward by using the front and rear limbs. These different types of moles with different excavation methods also relevant to the ground type. A naked mole-rat can open their mouth widely for a broad excavation area and have lip structures to prevent soil penetration [10-11]. Additionally, it has limbs for collecting and sweeping the excavated soil. It is possible to move more freely under the tunnel due to forming a wider borehole than the size of the body. This is possible owing to the high degree of freedom of the neck movement. These various excavation habits and\nConcept Design of a Novel Bio-Inspired Drilling System for Shallow Drilling Junseok Lee1, Christian Tirtawardhana1, Heung Woon Jang2, Jung-Wuk Hong2 and\nHyun Myung1* 1 School of Electrical Engineering, Korea Advanced Institute of Science and Technology (KAIST), Daejeon, 34141,\nSouth Korea (ljs630@kaist.ac.kr, christiant@kaist.ac.kr, hmyung@kaist.ac.kr) * Corresponding author 2 Department of Civil and Environmental Engineering, Korea Advanced Institute of Science and Technology (KAIST),\nDaejeon, 34141, South Korea (hwjang90@kaist.ac.kr, j.hong@kaist.ac.kr)\nAbstract: As the exhaustion of resources used as existing energy such as petroleum and coal are approaching, interest in planning and research for finding new energy sources is on the rise. In addition, the use of existing energy sources is one of the main causes of environmental destruction. Energy demand is steadily increasing, and it is urgently needed to find alternative energy sources due to exhaustion and limitation of existing energy sources. The use of drilling technology is inevitable for exploring these new energy sources. However, existing drilling equipment and technologies have limitations due to the colossal size and complicated processes, and also cause environmental damage resulting from drilling wide area. Therefore, in this paper, to overcome these limitations, the mechanism of a compact excavating robot that mimics the anatomy and habit of digging animals is proposed. The proposed bio-inspired mechanism simulates the excavating and debris removing process of a ground-digging mole using its teeth and forelimbs. Finally, the stress generated by the module during excavation is analyzed and the design durability is verified through the simulation. Keywords: New energy resources, drilling, mole, bio-inspired\nthe development of a suitable body structure evolved according to the environment in which they live.\nThe European mole has developed large forelimbs\nfor excavation and in search of prey in loamy soils where moisture is contained, while naked mole-rats develop additional front teeth to excavate harder soils with less moisture. European mole has a special bone structure for high excavation power of forefoot. The clue is scapula, related to shoulder bones. The human scapula has a round and flat shape, while the mole scapula has an elongated shape. By having the elongated shape of scapula, it has an extensive surface area, the connection of the greater number of muscles [12]. This biological structure can generate the high excavation force of the forelimbs than as compared with its body size.\nThe moles developed a specific structure according\nto the environment in which they survived. In this paper, shallow drilling with maximum ground strength about 3 MPa is aimed to excavate [15]. Therefore, a hybrid type excavation mechanism that mimics naked mole-rats\u2019 teeth for excavation as the drill bit and the debris removal of an excavated soil through the forelimbs of European mole is designed and proposed.\nFor excavation in soft ground, it is possible to use\nonly forelimbs but it is ideal to excavate solid ground using an additional structure like as teeth of naked mole-rat in case of about 3 MPa strength soil. Therefore, the excavation drill bit mimics the teeth and wide jaw movement range of the naked mole-rat. As a result, an expandable drill bit which can expand and contract is designed not by using an ordinary drill bit. The proposed expandable drill bit performs expansion/contraction function and rotation for drilling by using only one rotation motor. The developed drill bit consists of a body section for movement and a blade section for excavation.\nThe body section is composed with three parts in total. A fixed body, a rotating body, and a shaft. The shaft and the rotating body rotate simultaneously when the motor operates. The fixed body and the rotating body have screw patterns that mesh with each other like bolts and nuts. Consequentially, when the shaft and the rotating body are simultaneously rotated by the motor, the rotating body moves up and down by the screw pattern of the fixed body.\nTwo types of blades are connected to the rotating body. Among them, the expandable blades are also connected to the rack, and as the rotating body moves up and down, it rotates with a maximum angle of 90 degrees. When the rotating body descends as low as possible, the expandable blades rotate 90 degrees as clockwise and expansion. Conversely, in the case of a position where the rotating body moves up to the uppermost position, the blades rotate as 90 degrees counterclockwise and contraction.\nWhen drilling is progressed, the drill bit rotates in an expanded state like the widespread mouth of a naked mole-rat. Once the excavation is completed, the motor rotates in reverse, the screw patterns engage with each other, the rotating body moves up while the blades shrink. The lead time for expand and contract of the blades is designed to take about 2 seconds when the motor rotates at 300 rpm. Each blade is formed in a\nAuthorized licensed use limited to: Macquarie University. Downloaded on June 22,2020 at 19:52:00 UTC from IEEE Xplore. Restrictions apply.", + "1278\ntooth-like shape for efficient excavation. The expanded drill bit has a diameter of about 200 mm and the diameter of the excavated borehole will also form 200 mm. When excavation proceeds using the extended blades, it is advantageous to control the posture for the directional operation by securing the free space of the borehole.\nAfter forming the excavation hole of about 200 mm using the above-mentioned extended drill bit, the debris of the excavated soils is removed by using the mimicry structure of the European mole\u2019s scapula and forelimbs. The forelimbs are designed with the equivalent symmetrical structure on the top and bottom. Using the rotating motor attached to the upper part, the rack moves forward and backward through the movement of the crank shaft and the worm gear, and the force is distributed and transmitted to the lower part having the same structure by using connecting the belt. In addition, a linear actuator and a link structure are designed to mimic the mole scapula and anatomy on both sides. Synthetically, the forelimbs move forward through one rotary motor and two linear actuators and make the motion that gathering paws to the center, pushing arms to both sides and returning to the initial position. The removal of excavated soil proceeds through the repetition of this motion.\nFigure 6 shows the overall design combined expandable drill bit and forelimbs. Drill bit module\nkeeps advanced state while drilling, and when debris removal is started, it moves backward into the body, so that interference between the drill bit and forelimbs can be prevented.\nThe sequence of the whole drilling mechanism of the proposed robot is as follows. 1) Advance drill bit module, 2) Rotate while expanding blades, 3) Excavation, 4) Contraction blades and reverse drill bit module, 5) Move forward and folding of forelimbs, 6) Remove debris spreading and move backward forelimbs.\nAuthorized licensed use limited to: Macquarie University. Downloaded on June 22,2020 at 19:52:00 UTC from IEEE Xplore. Restrictions apply." + ] + }, + { + "image_filename": "designv11_80_0000948_ever.2019.8813582-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000948_ever.2019.8813582-Figure7-1.png", + "caption": "Fig. 7. Open-circuit flux distributions of the 12-slot/10-pole HESFPM machine when f1 is minimum and maximum.", + "texts": [ + " Nr is the rotor pole number. Therefore, the open-circuit DC winding flux-linkage and hence the induced voltage cycles per electric period Npe can be expressed by equation (4) [14]. (4) where LCM is the least common multiple. The open-circuit DC coil flux-linkage variation is caused by the variation of the equivalent air-gap reluctance when the rotor rotates. Taking the DC coil f1 as an example, the open-circuit DC coil f1 flux-linkage f1 varies with rotor electric position as shown in Fig. 6(a). As shown Fig. 7, when e=120o, f1 reaches its minimum value 9.39mWb while f1 reaches its maximum value 10.35mWb at e=300o. Consequently, the reluctance of DC winding\u2019s magnetic circuit under open-circuit condition is maximum at e=120o and minimum at e=300o due to the variation of the equivalent air-gap magnetic reluctance. The variation of the equivalent air-gap magnetic reluctance is essentially caused by the doubly salient structure of the HESFPM machine. VOLTAGE Skewing is widely used in electrical machines for reducing cogging torque, back-EMF harmonics and hence torque ripple" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000635_978-3-030-20377-1_1-Figure1.2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000635_978-3-030-20377-1_1-Figure1.2-1.png", + "caption": "Fig. 1.2 Incomplete contact which is nearly conformal", + "texts": [ + " Equally, if a shear force is applied so that shear tractions develop along the interface, the surface normal displacement of each body will be the same so that no change in relative surface gradient arises, and therefore the contact pressure distribution is unmodified. Contacts of this kind, where the normal problem induces no shear traction and shear tractions cause no modification of the normal pressure, regardless of the value of the coefficient of friction, are said to be \u2018uncoupled\u2019. Some contacts which are advancing in character do not permit representation of the bodies by half-planes. A good example is shown in Fig. 1.2 whichmay be thought of as the traditional civil engineering bearing bolt. The small clearance between the pin and hole in which it is journalled ( = R1 \u2212 R2 > 0 but R2) means that a formulation appropriate to an elastic disk and anti-disk is needed (Persson 1964). These contacts are also said to be \u2018nearly conformal\u2019. Awedge-shapedbodypressed into the surface of a secondbody is another example of an incomplete contact and ismentioned because surface elements of the contacting bodies will have to rotate as they pass into the contact patch under increasing normal load, with a combined rotation equal to the external wedge angle", + "1c, we have the extreme case of a line load on what may be an infinitely long strip\u2014the application of a finite normal load causes the contact to become much smaller, or to recede. In a problem of this kind where the basic geometry includes only a solitary length scale (here, the strip thickness, h) the contact \u2018snaps\u2019 to its final size upon application of an infinitesimal normal load. Increasing the force merely causes the state of stress and displacements to increase in proportion, but the separation points remain the same and, in all cases, p(s) \u223c \u221a s (if the bodies are elastically similar). Returning to the \u2018nearly conformal\u2019 contact illustrated in Fig. 1.2, we note that, when = 0 the application of a radial force to the pin causes the contact to \u2018snap\u2019 to an included angle of \u03c6 = 87.46\u25e6. This is also the angle to which an advancing contact extends when R > 0 and to which it smoothly recedes when \u2212R < 0 (Ciavarella et al. 2006). It will be apparent that, in all the contacts considered so far, the contact pressure adjacent to the contact edge can take one of only two values\u2014p(s) \u2192 0 as s \u2192 0 in the cases of advancing or receding contacts, and p(s) \u2192 \u221e as s \u2192 0 in the case of most stationary, complete contacts" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000303_978-981-13-3627-0-Figure1.4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000303_978-981-13-3627-0-Figure1.4-1.png", + "caption": "Fig. 1.4 Chip pick-up process through needle ejecting process: a the pick-up module with an amplified inset of the ejector needle, b three stages of the chip pick-up process. Reprinted from Ref. [89], with the permission of AIP Publishing", + "texts": [ + " The VCCT has been considered as an effective way to calculate the ERR of interfacial peeling [72, 74, 75, 83, 89, 90]. Xie and Biggers [100, 101] developed a kind of interface element called the fracture element with dummy nodes based on the VCCT, through which the ERR can be calculated simultaneously as the finite element analysis was performed. Besides, particular attention should be paid when the VCCT is applied in problems with bi-material interface cracks [97] or highly asymmetric cracks [102]. A description of the typical chip pick-up through needle ejecting process is shown in Fig. 1.4. The practical peeling-off module is also depicted. As shown in Fig. 1.4a, the pick-up process can be conveniently divided into three stages according to the operation flow. The three stages of contact-impact, peeling-off, and picking-up are featured by transient local impact, long quasi-static peeling-off, and convergent debonding, respectively. The chip peeling-off, which completes the transfer of the chip from the adhesive tape to the pick-up head, has to adapt to the flexibility and thinning of IC chips. The peeling-off stage occupies the longest time among these three stages and gets the own characteristic of quasi-static behavior" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003452_0954405420951101-Figure13-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003452_0954405420951101-Figure13-1.png", + "caption": "Figure 13. Finite element meshing model.", + "texts": [ + " While the numbers of teeth on gear increasing from 10 to 100, the reduction ratio of maximum root stresses increases from 2.99% to 3.84% with ISO model, but reduction ratio remains the same the modulus is 2, 4 or 6. A FEM model is calculated to verification the validity of theoretical model. HPSTC is calculated with the numbers of teeth on gear are 30. Applied a fixed constraint on the pinion hole and load a normal linear force 250N/mm on the HPSTC with different transaction curve. As is shown in Figure 13. There are 125,380 linear hexahedral elements of type C3D8R in Finite Element mesh division and a single tooth is refined for more reliable result. The results calculated by ABAQUS is shown in Figure 14. The local stress concentration due to constraint was ignored. The maximum root stress are 283.9MPa and 247MPa calculated with optimized transaction curve and standard transaction curve. The maximum root stress of optimized transaction curve and standard transaction curve are 191.4MPa and 172.8MPa calculated by AGMA model or 191" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000899_012082-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000899_012082-Figure1-1.png", + "caption": "Figure 1. The movement patterns of the end mill: a) rough milling of the root, b) fine milling, c) processing the root of the gear wheel.", + "texts": [ + " The more developed the CAM system is, the greater the number of options for setting the working paths it allows you to perform. Let\u2019s consider the most likely ways to handle toothed gear roots. Coordinates of work points can be calculated by all the listed software products in the basic version. ISTC-IETEM IOP Conf. Series: Materials Science and Engineering 570 (2019) 012082 IOP Publishing doi:10.1088/1757-899X/570/1/012082 The first is the transfer of rough milling which should eliminate the main allowance from the root, leaving it on the sides. The allowance is sampled by a spherical end or a keyway cutter (Figure 1a). Then comes the fine milling. It can be made with a modular end mill. This variant of processing assumes the shortest possible processing time of the wheel root, but the modular milling machine must be specially manufactured for parameters of the gear, its manufacturing cost may be comparable to the cost of manufacturing the worm gear-cutting milling machine. Another optional variant is the traversal by a spherical end mill (Figure 1 b). Unlike a modular disk or end mill, a spherical end mill is an inexpensive and affordable tool. The smaller the depth of cut, the better the surface of the side gears, but the longer the process is. In any case, to achieve an acceptable quality of the lateral surface of gears the processing time of toothed rims will be significant. In some cases it can reach 4-5 hours or more. The intermediate option among milling options discussed above (modular milling and spherical end milling) is, for example, Sandvik Coromant's offer in the form of \u201cpatented InvoMilling process\u201d [9] (figure 2)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001910_012002-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001910_012002-Figure3-1.png", + "caption": "Figure 3. Point normal vector and Point pose", + "texts": [ + "Index the two adjacent faces of the intersecting edge, find the other edges that intersect the intersecting plane and calculate the intersection. By analogy, the intersections are recursively reciprocated to the two sides until there is no intersecting edge or the intersecting edges coincide (circular path). Finally, a path point intersecting the section can be obtained. Analysis of the geometrical characteristics of the intersecting plane intersection point shows that the p0 point pose can be obtained from its normal vector and machining direction, as shown in figure 3. The normal vector of point p0 is determined by the characteristics of the surface. By indexing the data structures of the vertices p1 and p2 of the intersecting edges, the normal vectors e1 and e2 of the two points can be obtained respectively. Here, the position of p0 point between point p1 and point p2 is used as the weight, and the normal vector e0 of point p0 is calculated: 22110 ekeke (5) 21222111 // lllklllk (6) Where k1 is the weight of point p1, k2 is the weight of point p2, l1 is the distance between point p1 and point p0, and l2 is the distance between point p2 and point p0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000513_ab2747-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000513_ab2747-Figure2-1.png", + "caption": "Figure 2.The sketch of laser scanning; Themeshed 2D cross-section powder bedmodel andThe volumetric heat sourcemodel.", + "texts": [ + " Themain objective of this paper is to present the influence of the linear energy density (LED) of laser beam on themolten pool characters during SLMof PA6 powder. Simultaneously, the effects ofMarangoni convection and viscoelastic flowon themolten pool behavior were considered. Furthermore, in order to confirm the accuracy of the simulationmodel and optimize the process parameters in the SLMprocess, the relatively densities of the sintered parts were obtained via experiment andwere confirmed bymicrostructure and mechanical test. 2.1.Model andmesh Based onfinite elementmethod (FEM), a 2D cross-section geometricmodel was established. Figure 2 shows the sketch of laser scanning in the experiment andmodel. TheX axis represents the laser scan direction, and theY axis represents the depth of a series ofmolten pools. The powder bedwasmeshed according to the absorption properties of the laser bed. The gradient of the dropping temperature on the upper surface of the powder bed is larger, for increasing the efficency of simulation, the density of themesh isfinerwith rectangular cells in bigger size. As the depth increases, the rate of temperature drop becomes slower, and the correspondingmesh density becomes thicker with triangular cell in smaller size" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003853_ecai50035.2020.9223138-Figure7-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003853_ecai50035.2020.9223138-Figure7-1.png", + "caption": "Fig. 7. Assembly the core of the system with HEPA filter", + "texts": [ + " In order to reduce the mass of the device and its construction costs, when making the components by 3D printing, a honeycomb structure with a filling density of 40% was chosen as a solution for the internal structures of the components [7, 8, 9]. Thus, a design model was adopted based on a standard 120 mm fan. In figure 6 you can see how the air aspirate by the fan is passed through a HEPA filter, and then captured by a turbine in order to direct it to the equipment hose, ensuring to the air a laminar flow. In figure 7, the HEPA filter is assembled with an adapter element on the fan through a bayonet type system, and the assembly consisting of the turbine, fan and HEPA filter will form a sandwich, being easy to assemble / disassemble. The entire equipment is assembled in a case specially designed to be easily assembled (figure 8), with parts for self-guiding the components, but also for self-guiding the half-cases. At the same time, in the mechanical design of the product, it was taken into account to use materials easy to clean and sterilize, which would meet the norms imposed on the devices for medical use" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001835_s12206-019-1147-7-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001835_s12206-019-1147-7-Figure12-1.png", + "caption": "Fig. 12. Limit cycle generated by the CPG-LQR-RC combination for tracking a desired ZMP with multiple rolling speeds.", + "texts": [ + " Since the frequency F is the unique variable parameter in the system when dealing with multiple speeds, the period L of the desired output of the RC is the only variable in the RC design that needs to be tuned. Actually, all other parameters are not affected by changing the frequency of the rolling. Thus, to guarantee that the RC design is updated to fit any frequency F, the RC and the CPG are coupled such as L = 1/F. Here, L is the period of the desired output of the RC and F the frequency of the CPG. This method is tested on a gait starting at the walking frequency F1 = 0.78 Hz and continuing at F2 = 0.48 Hz. The Fig. 12 shows the convergence of the limit cycle of the controlled plant to the desired ZMP limit cycle. Coupling the frequency of the CPG to the RC frequency assure that the repetitive controller design is automatically adjusted to any frequency F of rolling. This enables the overall system to track a ZMP trajectory with multiple rolling speeds. In this paper, the use of repetitive control for ZMP trajectory tracking is addressed. The motivation behind using RC comes from the ability of the RC to assure perfect tracking of any periodic input" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002242_acemp-optim44294.2019.9007153-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002242_acemp-optim44294.2019.9007153-Figure1-1.png", + "caption": "Fig. 1: Construction of BLDC motor Linix 45ZWN24-40", + "texts": [ + " A mathematical equation for cogging torque is: Tcogg = \u22121 2 \u03c62\u03b4 dR d\u03d1r (1) where \u03c6\u03b4 describes PM flux crossing the air-gap, R describes total reluctance through which the flux passes. If the total reluctance R doesn\u2019t vary as the rotor rotates, the derivation in (1) is zero and also the cogging torque is equal to zero [21], [22], [7]. Cogging torque can be mathematically described also by Fourier series: Tcogg = \u221e\u2211 k=1 (Tksinsin(kLCM(Q, 2p)\u03d1r)+ +Tkcoscos(kLCM(Q, 2p)\u03d1r)) (2) Tksin and Tkcos describes amplitude of k-th harmonic component for sine or cosine part. LCM(Q,2p) represents least common multiple of number of slots Q and the number of poles 2p [7]. Fig. 1 shows construction of BLDC motor Linix 45ZWN2440, which is used for experimental verification. Motor is controlled by FOC. Motor has number of slots Q = 6 and number of poles 2p = 4. According to (2) cogging torque has 12 periods in one mechanical revolution. TABLE I: Parameters of used BLDC motor Linix 45ZWN24-40 Nominal power PN[W ] 40 Nominal current IN[A] 2,3 Nominal voltage UN[V ] 24 Nominap speed nN[rpm] 4000 Nominal torque TN[Nm] 0,097 Number of poles 2p[\u2212] 4 Stator resistance Rs[\u2126] 0,5255 D-axis inductance Ld[H] 0,000472 Q-axis inductance Lq[H] 0,000496 Moment of inertia J [kgm\u22122] 0,000009 Voltage constant Ke[ V s rad ] 0,013715 101 Authorized licensed use limited to: University of Exeter" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001763_aeat-04-2019-0087-Figure8-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001763_aeat-04-2019-0087-Figure8-1.png", + "caption": "Figure 8 The first variant of the edge controller", + "texts": [ + " The airplane prototype has not been not tested so far; but, two other experimental setups using the edge controller were tested and results are disclosed later in the article. The current prototype of the edge controller is the eight iteration designated as version 8 (v8.0). The first prototype was built in 2012 and went through numerous variants and experiments, including testing with different microcontroller families, MOSFET transistors, MOSFET driver circuits, LEDs, etc. The first iteration of the device can be seen mounted on an H-frame quadrotor experimental design. This testbed was developed in the early 2013 and is shown in Figure 8. It used two controllers mounted on the cross points of the fuselage. Each controller was managing two brushless direct current motors. This early design proved to be cumbersome and inadequate and all further versions from 2 to 8 aimed at controlling only a single UAVs distributed avionics paradigm Svetoslav Zabunov and Roumen Nedkov Aircraft Engineering and Aerospace Technology Volume 92 \u00b7 Number 2 \u00b7 2020 \u00b7 229\u2013236 BLDCmotor in one device. This first version of the module also relied on an obsolete by modern standards, but still capable for the day, 8-bit microcontroller manufactured in a large dual inline package clearly visible in Figure 8. As noticed earlier, all further version of the device used modern ARM-core microcontrollers, namely, variants of the Cortex-M01 core microcontrollers from NXP. This selection was done for the following reasons: the extremely high performance to price ratio; high EM noise immunity; inherent resistance to ionizing radiation damage up to 50 krad(Si) (Leite et al., 2017); low power consumption; versatility; adequate, comfortable and free programming environment; and accessible and inexpensive function-rich programmer and debugger" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001144_s1068798x19090168-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001144_s1068798x19090168-Figure1-1.png", + "caption": "Fig. 1. Geometry of transceiver housing with built-in cooling channel on the printed-circuit side (a) and on the opposite side (b): (1) location of basic heat-liberating elements; (2) cooling channel; (3, 5) input and output of cooling channel; (4) far side of housing, where the cooling channel turns; (6) reinforcing rib.", + "texts": [ + " At present, the production of components with built-in 78 cooling channels by 3D printing of metals is under active development [14, 15]. In the present work, we describe a new housing for the transceiver of an active phased array antenna with a built-in cooling channel. The channel passes along the housing wall, and its geometry meets three basic requirements: effective cooling of the active elements; compatibility with selective laser melting (SLM) [16, 17]; and easy removal of the powder remaining in the channels after three-dimensional printing. The housing geometry in Fig. 1 is developed on the basis of preliminary hydrodynamic calculations. 5 The housing includes a metal box with connectors under the printed circuit and the transceiver components and an internal cooling channel running along the opposite side. The cooling channel first runs from the housing input (3 in Fig. 1) directly to the opposite side of the housing (4 in Fig. 1) and specifically to the position of the active UHF elements, where the heat liberation is most intense, and then runs, in coiled form, over the whole housing surface to the output (5 in Fig. 1), which is adjacent to the input. The cooling channel is hexahedral (width 25 mm, height 4 mm) so as to facilitate removal of the residual powder (Fig. 2a). The thickness of the channel wall is 1 mm; the lower wall is actually the wall of the transceiver housing. Within the channel, additional barriers (small columns) located under the UHF elements, where heat liberation is greatest, intensify the heat transfer from the surface to the cooling liquid (Fig. 2b). To ensure the required surface roughness, the transceiver housing is machined after manufacture" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000090_052079-Figure12-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000090_052079-Figure12-1.png", + "caption": "Figure 12. An example of a free-form grillage made from the identical beams.", + "texts": [ + " Figure 11 shows a beam made from a rectangular tube with an added flat bar that widens the base and plates that increase the beam support surface area. The choice of the cross-section is justified by the creation of conditions for supporting the elements that fill the grillage boxes. The holes facilitate the attachment of adjoining beams in alternate places. The beams according to figure 11 make it possible to construct grillages with a free-form configuration based on a design grid with module m (figure 12). WMCAUS 2018 IOP Conf. Series: Materials Science and Engineering 471 (2019) 052079 IOP Publishing doi:10.1088/1757-899X/471/5/052079 WMCAUS 2018 IOP Conf. Series: Materials Science and Engineering 471 (2019) 052079 IOP Publishing doi:10.1088/1757-899X/471/5/052079 An issue requiring a separate analysis is the method of making the grillage from short beams, including the order of assembly of individual elements and the solution of supporting the structure for the time of assembly (figure 13). Since none of the beams is supported solely on the walls or columns, each requires a temporary support" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001231_demped.2019.8864909-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001231_demped.2019.8864909-Figure2-1.png", + "caption": "Fig. 2: A portion of an eight-pole synchronous reluctance motor with three flux-barriers per pole.", + "texts": [ + " In the prototype machine new windings were installed and one phase winding was center tapped with multiple taps to vary the amount of short circuited turns during fault simulation. A SynRM rotor, with three flux-barriers per pole, has been optimized to achieve good anisotropy and low torque ripple both in healthy and faulty conditions. Some winding arrangements have been investigated, keeping the optimal geometry to be fixed, to the aim of improving the machine performance in faulty conditions. A sketch of a 8-pole synchronous rotor with three flux-barriers per pole is reported in Fig. 2. All the analyses are performed by means of finite element (FE). The complete details of the optimization will be provided in the full paper. After that, further optimizations of rotor and stator geometry, with different objectives and parameters, are carried out to find the solution that achieves the best trade-off between 978-1-7281-1832-1/19/$31.00 \u00a92019 IEEE 251 the healthy and faulty performances. A prototype of the motor is under construction and will be tested in several working conditions to validate the FE analysis results" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001194_1350650119877059-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001194_1350650119877059-Figure1-1.png", + "caption": "Figure 1. Geometry of ball bearing and contacting elastic solids.50", + "texts": [ + " The present exploration is done for elliptical contact lubricated with micropolar fluids following the work of Hamrock and Dowson. The problem model for an isothermal steady-state EHL comprises the Reynolds equation, the force balance equation, and the film thickness equation with elastic deformation of contacting solids. Two ellipsoids can represent the undeformed geometry of contacting solids. The two contacting solids brought together initially make contact at a point when the load is zero, called point contact as shown in Figure 1. Geometric parameters of contacting solids are also shown in the figure. Curvatures of solids described in x and z directions are given as 1 Rx \u00bc 1 rax \u00fe 1 rbx and 1 Rz \u00bc 1 raz \u00fe 1 rbz For ball to inner race contact Rx \u00bc d \u00f0de d cos \u00de 2de Rz \u00bc rid 2ri d \u00f01\u00de and for ball to outer race contact Rx \u00bc d \u00f0de \u00fe d cos \u00de 2de Rz \u00bc rod 2ro d \u00f02\u00de When these two contacting elastic solids are pressed together under externally applied load, the point contact extends and forms an elliptical contact. This contact area is described by the elliptical parameter k\u00bc a/b, where b and a are the semiminor and semimajor axes, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001376_j.mechmachtheory.2019.103672-Figure15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001376_j.mechmachtheory.2019.103672-Figure15-1.png", + "caption": "Fig. 15. Forces transmitted by G4 and G5.", + "texts": [ + " In order to assess the nature of the forces transmitted to the chassis of the testbench or vehicle, close attention should be paid to two of the gears, for instance, G4 and G5. When the sun turns clockwise, these gears transmit power and the forces are transmitted to the chassis; however, when the sun turns counter-clockwise, the freewheel inside G4 (negative freewheel) is decoupled and the forces become negligible. Because of this, these forces are intermittent and consequently a source of vibrations (Fig. 15). In a similar way to the previous case, a possible solution to reduce the vibrations is to use a flywheel at the output shaft. In this way, the output shaft rotation will be smoother, which leads to a reduction of the vibrations. However, the acceleration of the output is affected and, again, a compromise must be reached. Another more interesting solution is to use two epicyclic trains working in parallel. This solution is characterised because the whole transmission consists of two epicyclic trains working together, driven by two four-bar mechanism which provide motions out of phase 90\u00b0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002178_978-3-030-38077-9_91-Figure14-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002178_978-3-030-38077-9_91-Figure14-1.png", + "caption": "Fig. 14. Leila freight frame bogie and 4L bogie evolution", + "texts": [ + " For the rapid freight bogie developments in European rails, there are two outstanding designs of lateral swung bolsters and rubber cushions under central bowl, by which the two issues have to be considered carefully as follows: (1) Side-bearing friction instability. Similar to the Y37 frame freight bogie, the suspending rings with articulated joints, as shown in Fig. 13, can be used so as to avoid the abrasion impacts. But the dynamical simulation analyses indicate that the side-bearing friction instability problem will take place on the lateral swung bolsters when speeding up 160/250 km/h. (2) Vibration fatigue safety. For Leila frame freight bogie, as shown in Fig. 14a, the rubber cushion design under central bowl doesn\u2019t need the traction-rod devices but the technical difficulties will be solved costly thereby. So the evolution design of 4L bogie is trying to use the special pyramid frame instead of rubber cushion, as shown in Fig. 14b, which is impossible to guarantee the vibration fatigue safety in both tare and laden conditions. As the three particular demands stated above, the light weight carbody designs for light freight wagons are more difficult than the ones for conventional freight wagons. If the lower dynamical interactions of wore wheel-rail contact can not be achieved in the tare loopback, the light weight carbody vibration will become stronger and stronger, by which the construction velocity (or design speed) will be decreased thereby", + " Simulation result of carbody yaw angle in R25m 0 10 20 30 40 -6 -4 -2 0 2 4 6 antikink-MC2 antikink-M antikink-T antikink-MC1 no-antikink-MC2 no-antikink-M no-antikink-T no-antikink-MC1 ya w a ng le / \u00b0 time /s Fig. 12. Simulation result of carbody yaw angle in brake fault condition 0 10 20 30 -6 -4 -2 0 2 4 6 antikink-MC2 antikink-M antikink-T antikink-MC1 no-antikink-MC2 no-antikink-M no-antikink-T no-antikink-MC1ya w a ng le / \u00b0 time /s Fig. 13. Simulation result of carbody yaw angle in push rescue condition 0 10 20 30 -6 -4 -2 0 2 4 6 antikink-MC2 antikink-M antikink-T antikink-MC1 no-antikink-MC2 no-antikink-M no-antikink-T no-antikink-MC1 ya w a ng le / \u00b0 time/s Fig. 14. Simulation result of carbody yaw angle in traction fault condition Study on the Mechanism and Influencing Factors 559 3.2 Experimental Verification As shown in Fig. 15, a displacement laser sensor is installed between the bogie and the carbody to actually measure the maximum displacement between the bogie and the carbody when a tram passes the R25m curve. The test results (Fig. 16) show that, the anti-kink system can keep the yaw angle of the carbody consistent. Comparing the experimental results with the simulation results, it can be found that the experimental results and the simulation results tend to be consistent", + " 15(c)), the redistribution of forces happens to the maximum level, and the net change in lateral force is positive at mid-range speeds and negative at very high speeds. Unlike the previous case, the operating region includes linear, transition and saturation regions also. The gain in cornering performance at mid speeds is like in the previous case, in which the operating region is linear. With too much redistribution, the front Fig. 11. Tire workload difference Fig. 12. TW on inner and outer tire Fig. 13. Yaw rate Fig. 14. Lateral acceleration Development of Velocity Dependent Front Wheel Angle Relation 1329 outer tire is operating beyond the peak value, and the front inner tire is also not operating near its peak, so the net lateral force difference at very high speeds is negative. To evaluate the transient handling response, a 4 parameters analysis was carried out by Mimuro [9]. The evaluation was carried out by giving a pulse steer input at different vehicle speeds and getting yaw rate and lateral acceleration response" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000257_2019001-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000257_2019001-Figure1-1.png", + "caption": "Fig. 1. Three-dimensional diagram of tripod sliding universal joint and lubricant film.", + "texts": [ + " Energy conservation equation Conservation of energy means that the sum of mechanical energy and all internal energy in a system is equal to the total energy of the system. According to the energy conservation equation, control body energy variation equals the sum of influent energy, surface force work, volume force work and influent heat: DE Dt \u00bc DQ Dt \u00feDW Dt \u00f03\u00de In formula (3), E is the internal energy of the system, Q is the heat transferred into the system from the external environment and W is the external work on the system. The Pro/E modeling system is used to build the model of the tripod sliding universal joint and lubricant film [12], as shown in Figure 1. As the lubricant film between the tripod sleeve and sliding pin in the tripod sliding universal joint is very thin, the ratio between the lubricant film length and thin wall dimension is very large, which makes meshing difficult. Therefore, an unstructuredmesh is created using the octree algorithm in ICEM (shown in Fig. 2a), which has superior robustness. Although unstructured mesh facilitates complex area meshing, when the unstructured mesh is imported into FLUENT for solving, the mesh may result in a significant expansion ratio that could affect calculation precision" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001288_j.matpr.2019.08.229-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001288_j.matpr.2019.08.229-Figure10-1.png", + "caption": "Fig. 10. Results of longitudinal stress contours (a) Unidirectional epoxy carbon prepreg (230 GPa), (b) Unidirectional epoxy carbon Wet (230 GPa), (c) Unidirectional E-glass epoxy, (d) E-glass epoxy wet, and (e) Epoxy resign.", + "texts": [], + "surrounding_texts": [ + "Fig. 9. Results of hoop stress contours (a) Unidirectional epoxy carbon prepreg (230 GPa), (b) Unidirectional epoxy carbonWet (230 GPa), (c) Unidirectional E-glass epoxy, (d) E-glass epoxy wet, and (e) Epoxy resign.\n5. Analysis on composite pressure cylinders\nComposite pressure cylinders are made with UD Epoxy Carbon (230GPa) Prepreg and the results are shown in Tables 1 and 2 and Figs. 7\u201314. The results shows that, composite material weight is\nPlease cite this article as: K. Rajendra Prasad and C. Syamsundar, Theoretical applications, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.201\nless as compared with conventional materials like aluminum alloy. Among different composite materials, unidirectional epoxy carbon (230 GPa) Prepreg composites will shows less stresses and deformation. It is concluded that, composite material with same thickness have less weight as compare to aluminum alloy.\nand FE analysis of epoxy composite pressure cylinder used for aerospace 9.08.229", + "Fig. 12. Results of equivalent stress contours of alloys (a) Aluminum (Al), (b) Copper (Cu), and (c) Structural steel materials.\nFig. 13. Results of hoop stress contours of alloys (a) Aluminum (Al), (b) Copper (Cu), and (c) Structural steel materials.\nTable 3 Comparison of unidirectional epoxy carbon and aluminum alloy cylinders.\nS.No. Material Density (kg/m3) Thickness (mm) Diameter (mm) Volume (m3) Mass (kg)\n1. UD epoxy carbon (230 GPa) Prepreg 1490 101.54 1879 7.0248 10,467 2. Aluminum Alloy 2712 101.54 1879 7.1751 19,459\n6. Conclusions\nIt is concluded that, composite materials will have high strength by comparing with all materials and alloys. For 14 MPa, internal pressures of CNG Auto applications, Composite materials are the best choice. It is also found that UD Epoxy Carbon (230 GPa) Prepreg will have less deformation, as compared to other composites.\nReferences\n[1] P.Y.T. Abkov, E.B. Summers, Lay-up optimization of multilayered anisotropic cylinders based on 3-D elasticity solution, Comput. Struct. 84 (2006) 374\u2013384.\nPlease cite this article as: K. Rajendra Prasad and C. Syamsundar, Theoretical applications, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.201\n[2] Lei Zu, Sotiris Koussios, Adriaan Beukers, Design of filament-wound isotensoid pressure cylinders with unequal polar opening, J. Compos. Struct. 92 (2009) 2307\u20132313. [3] Shafique M.A. Khan, Stress distributions in a horizontal pressure cylinder and the saddle supports, Int. J. Pressure Cylinders Piping 87 (2008) 239\u2013244. [4] A. Alibeigloo, Static and vibration analysis of axi-symmetric angle-ply laminated cylindrical shell using state space differential quadrature method, Int. J. Pressure Cylinders Piping 86 (2008) 738\u2013747. [5] A. Hocine, D. Chapelle, M.L. Boubakar, Experimental and analytical investigation of cylindrical parts of a metallic cylinder reinforced by filament winding while submitted to internal pressure, Int. J. Pressure Cylinders Piping 86 (2009) 649\u2013 655. [6] K.V.J. Raoand, K. Narayana Rao, Design and analysis of filament wound composite pressure cylinder with integrated end domes, Defence Sci. J. 59 (1) (2009) 73\u201381. [7] M.W. Hyer, Stress Analysis of Fiber Reinforced Composite Materials, The McGraw-Hill Companies Inc., USA, 1998.\nand FE analysis of epoxy composite pressure cylinder used for aerospace 9.08.229" + ] + }, + { + "image_filename": "designv11_80_0000925_b978-0-12-816634-5.00003-0-Figure3.15-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000925_b978-0-12-816634-5.00003-0-Figure3.15-1.png", + "caption": "Figure 3.15 Images of the FA-18 rudder antirotation bracket LC repair: (A) pin area damage before repair, (B) after depositing over the grind out area, and (C) after machining [39].", + "texts": [ + " Other components that can be repaired include housings, bearings, casing flanges, seals, landing gears, etc. Fig. 3.14 shows a DMD-repaired die-casting tool [38]. This H13 tool was extensively worn and the core cracked during an aluminum diecasting process. The tool was premachined to remove damaged material and then rebuilt with 300 Maraging steel using the DMD process. In some cases, AM repair may not be cost effective, however, it offers a significant reduction in lead time and may be critical for defense applications. An example of such a benefit is depicted in Fig. 3.15 [39]. An F/A18 rudder antirotation bracket has been repaired using a laser-based DED process. The original bracket material was 17-4 PH stainless steel with a hardness of HRC 35 38. In order to match this hardness, a mixture of gas-atomized stainless steel SS316 and SS420 powders was used to form deposited layers over the grind out area using a high-power laser and a six-axis robotic equipment with a deposition head. Even though the replacement part cost for this application was not high, the lead time was very long (18 months) and AM repair can be achieved in weeks, making it a useful tool for the end user" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002099_icar46387.2019.8981581-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002099_icar46387.2019.8981581-Figure6-1.png", + "caption": "Fig. 6: Desired contact between the structural-frame and the panel", + "texts": [ + " Figure 4 shows the overall schema to command a position-controlled robot using our Multi-Surface admittance control approach. The case study in this paper is the robotic assembly of two aircraft parts shown in Figure 5. The first part is a fixed panel (3 meters high and 4 meters wide) and the second is a structural-frame, over 2 meters high and weighs around 20kg, handled by the robot. There are six contact surfaces between the two parts. After a successful assembly operation, there must be no clearance (< 0.3mm) between the parts and the contact forces must be uniformly distributed. Figure 6 illustrates the desired contact between the parts as well as the position of the force sensor. All the forces applied to the structural-frame are measured related to the sensor\u2019s coordinate system Sxyz. Due to the poor accuracy of robots, and the uncertainties of the position and orientation of the panels, the robot\u2019s trajectory must be adjusted during assembling processes. Figure 7 presents two possible contact scenarios: when the first contact happens with the upper surface of the structural-frame, and the second when it happens with the bottom surface" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000340_s11665-019-04078-z-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000340_s11665-019-04078-z-Figure1-1.png", + "caption": "Fig. 1 Schematic diagram of: (a) organics cladding powder and (b) new laser side feeding device", + "texts": [ + " Here, we apply a new organics cladding pretreatment technique and a new laser side feeding device to provide a protective atmosphere for molten pool, aiming at refining the grains and alleviating defects which could improve mechanical properties of austenitic steel coating fabricated by the laser remanufacturing technology. The austenitic steel powder (18-19% Cr, 8-9% Ni, 0.12- 0.15% C, 0.08-0.1% B, 0.9-1.0% Si, 1.5-2%Mn, 0.1-0.2% V and Fe surplus) with a size of 150 250 mesh was used. The powder was first coated by the colophony with 100 nm thickness and then covered by the fecula with 200 nm thickness (Fig. 1a). For the preparation of the colophony coating, rosin was dissolved in alcohol, and then the powder was added to the solution of colophony\u2013alcohol, stirred and heated. After the alcohol evaporated, rosin formed a layer of uniform rosin coating on the surface of the metal powders. The process of preparing fecula coating is similar to that of preparing rosin coating: fecula was dissolved in cold water, and heated and stirred to form a uniform aqueous starch solution. Then the powders covered with the rosin coating was added to the solution, and continued heating and stirring at the same time", + " Yong Chen, Liangbin Hu, and Changjun Qiu, School of Mechanical Engineering, University of South China, Hengyang 421001, China. Contact e-mail: qiuchangjun@hotmail.com. JMEPEG \u00a9ASM International https://doi.org/10.1007/s11665-019-04078-z 1059-9495/$19.00 Journal of Materials Engineering and Performance The Q235 steel (0.35% Si, 1.4%Mn, 0.22% C and Fe surplus) was adopted as a substrate, which was machined into plates with a dimension of 360 mm9120 mm924 mm. The coated powders were fed by the new synchronous lateral powder feeding device (Fig. 1b). The parameters of laser forming are as follows: laser power was 2.3 kW, the size of laser beam spot with ellipsoidal shape was 3 mm94 mm, laser beam spot scanning speed was 6 mm/s, the powder feeding rate was 6.7 g/min, and the overlap rate was 50%. The pulse heating inert gas fusion-infrared absorption method was applied to determine the oxygen content. Tensile experiments were conducted under the rate of extension of 0.2 mm/min at ambient temperature. Micro-hardness of specimens was measured using a digital tester under an indenter load of 1" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001428_012115-Figure4-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001428_012115-Figure4-1.png", + "caption": "Figure 4. Finite-element model of rolling cylindrical billets", + "texts": [ + " Thus, the obtained value of the limiting capture angle \u03b1 in the range of 2 \u00b0 -8 \u00b0. The maximum value of the absolute reduction depends on the friction coefficient and the workpiece diameter. To determine its optimal value, let us perform a simulation of the process of cylindrical billets transverse running with flat plates. To determine the stress-strain state in the deformation zone and residual stresses in the straightened parts, a finite-element model in the form of a cylinder and two plates was built in the Ansys workbench program (Fig. 4). The following simulation parameters were adopted: a cylinder with a diameter of 10 mm, a length of 100 mm from steel St45 (yield strength \u03c3\u0442 = 360 MPa and an elastic modulus of E = 2 * 105 MPa); working plates with dimensions of 5x110x110 mm are considered absolutely rigid. Simulation parameters: finite element form \u2013 hexahedron, thickening of 7030 elements, 30620 knots; coefficient of friction between the workpiece and the plates \u03bc = 0.15; ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012115 IOP Publishing doi:10" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002301_icisct47635.2019.9011865-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002301_icisct47635.2019.9011865-Figure3-1.png", + "caption": "Fig 3. Kinematic pairs degrees.", + "texts": [ + " The analysis of the results is carried out - the positions of the parts, velocities, accelerations, reactive forces and torques, driving forces and their moments are determined. Dynamic Simulation module gives an extensive set of kinematic pairs (Fig 2). It provides both standard types of kinematic pairs and various special types that describe the operation of gears and worm-gears with moving and fixed axles, belt and chain gears, cam mechanisms, ratchet and pinion gears, and set elastic ties and three-dimensional contact between the bodies. Kinematic pairs are set by two methods. The first method is the creation of kinematic pairs by setting the degrees of freedom (Fig 3). To do this, select the desired type of kinematic pair from the menu, select the parts that form the kinematic pair and then connect them together, defining edges, axes, planes, points, etc. of interacting components. After the end of operation, the parts are automatically set to the specified location [5]. The second method is to use assembly dependencies created in the context of the Autodesk Inventor assembly (Figure 4). To do this, select two parts that form a kinematic pair, and activate the existing assembly dependencies For the model to work, it is necessary to set to its nodes the appropriate movements" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003311_speedam48782.2020.9161923-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003311_speedam48782.2020.9161923-Figure3-1.png", + "caption": "Fig 3. Calculated magnetic flux density distribution after short-circuit from motor no-load condition at t = 4.8 ms (JMAG Designer) (compare Fig. 2)", + "texts": [ + " The initial values, the used machine parameters and the peak values of the short-circuit current in phase U iU,peak and the electromagnetic torque me,peak are summarized in Table II for all considered load cases of the SynRM. By using the synchronous inductances obtained from load operation, the analytically calculated peak short-circuit current is 30% lower and the absolute value of the peak short-circuit torque is 35% lower than the numerically calculated values. The deviations are caused by high iron saturation in the q-axis. Fig. 3 shows the magnetic flux density plot at t = 4.8 ms, when the maximum short-circuit current occurs. The magnetic flux density in the stator tooth, the stator yoke and the rotor iron bridges are quite high. The magnetic field is nearly aligned in the q-axis, so that the maximum shortcircuit current is almost a pure q-axis current. This coincides with the results of the analytical calculations in (5) and (6). Assuming a saturated q-axis inductance Lq, the analytical calculations agree much better to the numerical calculations (Fig" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000132_j.jfluidstructs.2019.02.019-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000132_j.jfluidstructs.2019.02.019-Figure3-1.png", + "caption": "Fig. 3. Principle of the cantilever method.", + "texts": [ + " We used a high-speed camera at a rate of 2000 images per second to record the movement of the filament. A sodium lamp was chosen as the only light source since its wavelength is similar to the thickness of film such that the optical light reflecting on upper and lower surfaces forms interference fringes that allow the visualization of the flow field. The filaments we used were made of spandex, polyester and silk with several different diameters. The diameter of the filament was measured by using a scanning electron microscope as shown typically for spandex filaments in Fig. 2. As shown in Fig. 3, a cantilever method was applied to obtain the bending stiffness of the filament that was estimated as follows (Dadeppo and Schmidt, 1971): B = mLgL3 cos( \u03b8 2 )/8 tan \u03b8 (1) where mL denotes the linear density of the filament, g denotes the acceleration due to gravity, and \u03b8 denotes the inclination to the horizontal of the straight line OA. A few important properties of the filaments are listed in Table 1. Zhang\u2019s experiment (Zhang et al., 2000) is listed at the same time as a comparison. The thickness of the soap film in parallel part hp was obtained as follows: hp = Q/(Lp \u00d7 U) (2) where Q denotes the quantity of flow per second, Lp denotes the width of film in parallel part, and U denotes inflow velocity" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001428_012115-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001428_012115-Figure2-1.png", + "caption": "Figure 2. Scheme of the workpiece capture by the flat plates: 1 \u2013 movable plate, 2 - blank, 3 - fixed plate, 4 \u2013 stops", + "texts": [ + " To implement this process of running it is necessary to determine its main parameters: the capture angle of the workpiece, absolute compression, stress state in the deformation zone, residual stresses after running. To carry out the process of running the blank with flat plates, it is necessary to create certain conditions. Moreover, it is necessary to consider separately the conditions for the unsteady running process \u2013 for ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012115 IOP Publishing doi:10.1088/1757-899X/632/1/012115 the initial moment when the workpiece is only supplied to the plates (Fig. 2) and for the steady-state process, when the workpiece is already drawn into the plates (Fig. 3). Consider the scheme of the workpiece capture by the flat plates (see. Fig. 2). In the lead-in part, the elevation angle \u03b11 is the main geometrical parameter. In the case of an unsteady running process at the moment when the workpiece contacts the plates, the latter will act on it in the form of N forces directed normally to the shafts surface at the points where the workpiece contacts the plates A, and friction forces F directed tangentially, as shown in fig. 2 In order to identify the effect of the indicated forces N and F on the conditions for gripping the workpiece with plates, we project them onto the horizontal axis XX (along the running direction) and onto the vertical axis YY. To carry out the rolling process (axis XX), the horizontal force Fx must be greater than the force Nx, then we have: 11 sincosF N 11 sincosN N 1tg arctg1 , (1) where \u03bc \u2013 coefficient of friction between the workpiece and the plates. The value of the friction coefficient for a steel-steel pair during cold running is within the range from 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003984_012029-Figure10-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003984_012029-Figure10-1.png", + "caption": "Figure 10 Fourth-order mode", + "texts": [], + "surrounding_texts": [ + "The modal analysis is executed in the modal analysis module of ANSYS Workbench software, and then the geometric model of the brake disc is imported, the material properties are set, the brake disc is meshed, and the modal analysis result of the brake disc is finally obtained. When analyzing the modal results, select the first ten vibration frequency value and mode description of the brake disc as shown in the following figure (Figure 6-15): ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 ICAMIM 2020 Journal of Physics: Conference Series 1653 (2020) 012029 IOP Publishing doi:10.1088/1742-6596/1653/1/012029 4 1226.4 Around the center, it deforms like petals 5 1230.2 Symmetrically deformed on both sides along the Y axis 6 1806.9 Outer ring deformation 7 2542.3 Wavy deformation 8 2601 Symmetrically deformed along the friction block 9 2706.1 Curved toward the center in a petal shape 10 2708.5 Bending around the Z axis and deforming along the Y axis From the above tenth order frequency distribution, it can be seen that the lowest order frequency and the first order frequency of the brake disc are about 630Hz. Due to different brake structures, the squeal frequency of this disc brake is about 2100Hz. The modal analysis result shows 6 The first and seventh-order modal frequencies are closer to the squeal frequency, and there is a greater possibility of screaming." + ] + }, + { + "image_filename": "designv11_80_0001428_012115-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001428_012115-Figure3-1.png", + "caption": "Figure 3. Endeavors scheme during transverse running in flat plates: 1 - movable plate, 2 - blank, 3 - fixed plate, rk - radius of the elastic core", + "texts": [ + " To carry out the process of running the blank with flat plates, it is necessary to create certain conditions. Moreover, it is necessary to consider separately the conditions for the unsteady running process \u2013 for ICI2AE 2019 IOP Conf. Series: Materials Science and Engineering 632 (2019) 012115 IOP Publishing doi:10.1088/1757-899X/632/1/012115 the initial moment when the workpiece is only supplied to the plates (Fig. 2) and for the steady-state process, when the workpiece is already drawn into the plates (Fig. 3). Consider the scheme of the workpiece capture by the flat plates (see. Fig. 2). In the lead-in part, the elevation angle \u03b11 is the main geometrical parameter. In the case of an unsteady running process at the moment when the workpiece contacts the plates, the latter will act on it in the form of N forces directed normally to the shafts surface at the points where the workpiece contacts the plates A, and friction forces F directed tangentially, as shown in fig. 2 In order to identify the effect of the indicated forces N and F on the conditions for gripping the workpiece with plates, we project them onto the horizontal axis XX (along the running direction) and onto the vertical axis YY", + " To carry out the rolling process (axis XX), the horizontal force Fx must be greater than the force Nx, then we have: 11 sincosF N 11 sincosN N 1tg arctg1 , (1) where \u03bc \u2013 coefficient of friction between the workpiece and the plates. The value of the friction coefficient for a steel-steel pair during cold running is within the range from 0.03 to 0.15 [12-15]. According to the formula (1), we obtain the value of the angle \u03b11, which is in the range from 2 \u00b0 to 8 \u00b0. Next, we consider the scheme of transverse running with flat plates at the steady-state process, when the workpiece is already drawn into the space between the plates [16]. As Figure 3 demonstrates, the top plate movement direction is perpendicular to the rotation axis of the cylindrical body being rolled. The distance between the plates is less than the original diameter of the cylinder by 2y - this is the value of the absolute compression. The endeavors are used towards the workpiece in the direction of the central axis that are directed normally to the contact pad. The resultant of these efforts P will be considered as applied in the middle of the segment corresponding to the zone where the workpiece contact the plates" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001638_978-981-15-1263-6_2-Figure1.11-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001638_978-981-15-1263-6_2-Figure1.11-1.png", + "caption": "Fig. 1.11 AM possibility for reducing piston weight through lattice integration (Figure adapted from Reyes, Belmonte et al. (Angel et al. 2015))", + "texts": [ + " 1.10) with open internal cavities for routing wiring harnesses. The arms have no external penetrations and are designed to be neutrally buoyant and, due to the simplicity of the design enabled by AM, can be disassembled or assembled in under an hour. As another example, a simulation assessment of lattice structures for use in engine components was conducted by the University of Bath. The results illustrated the potential for integrating a lattice into the center section of a piston, as shown in Fig. 1.11. The simulation prediction results suggest it may be possible to reduce piston mass by 9% (Angel et al. 2015) with little to no change in structural integrity. This also would allow reducing the connecting rod and crankshaft mass due to lower operating loads imposed by the pistons. Such system impacts need to be included in revised cost models. Recently, an AM optimization and opportunity study of a Delphi-based diesel fuel pump design was conducted by the AM consultant Econolyst (now the Strategic 1 Opportunities for Lighter Weight and Lower Total Cost Component \u2026 15 Consulting Team at Stratasys) in association with Loughborough University (Benatmane 2010)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001521_042047-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001521_042047-Figure3-1.png", + "caption": "Figure 3. Four gating system of impeller. Top gating system(a), bottom gating system(b), liner-sharp side gating system(c) and cross-sharp side gating system(d).", + "texts": [ + " Before the practice of manufacture, AnyCasting is applied for design of gating system. The parameters of process are chosen depend on reality fabricate condition of workshop. Thus the results of production will only due to the management of different gating system. The 3D model is generated according to the setting of size in figure 2. Then the model that including four types of gating system are created. Top gating system, bottom gating system and 2 types of side gating system are added to the impeller as Figure 3. ICEEMS 2019 IOP Conf. Series: Earth and Environmental Science 332 (2019) 042047 IOP Publishing doi:10.1088/1755-1315/332/4/042047 2.2.Simulation of gating system The software Anycasting bears the simulation. The mesh are created in the preprocessing module called AnyPRE. Taking the top gating system as the example, the account of flexible finite elements is 103488 in Figure 4. Setting the case as the filling analysis and solidification analysis, then the results of preprocessing are saved for later treatments" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003038_012194-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003038_012194-Figure1-1.png", + "caption": "Figure 1. Model for mountain tractor suspension implements", + "texts": [ + " In order to control the suspension implement position posture and adjust the depth of the mountain tractor implement, the three degree of freedom of the mountain tractor suspension implement is presented in this paper. A tractor model with three degree of freedom agricultural implement suspension mechanism is set. The model can be used to achieve the control of the position and posture of the tractor implement, and adjust the tillage depth of the agricultural implement and the contour of the slope[9-10]. The suspension implement model is shown in Fig1. Here the three degree of freedom suspension device is composed of a left telescopic pulling rod, a right telescopic pulling rod, a pull-down rod, and an agricultural implement attachment frame. It can be seen from the two above figures, the left and right telescopic rods are devices with moving pairs, which can be extended or shortened. One end of the left telescopic rod and the right telescopic rod is hinged with the tractor body through a ball pair, and the other end is hinged with the ball pair to the frame of the tractor implement" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002792_s12046-020-01366-6-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002792_s12046-020-01366-6-Figure6-1.png", + "caption": "Figure 6. Watt mechanism with slider (2).", + "texts": [], + "surrounding_texts": [ + "ADAMS allows export of data in the form of an Excel Sheet. The data contains many data points of the graph at equal intervals of time. Thus, we have two data sets for the two curves. Now, we need to compare these data sets in order to quantify the difference between the two curves. Some statistical measures which can be used to compare the two data sets are mean deviation (obtained by taking the mean of differences in output of mechanisms with and without clearance), RMS deviation, 90% points tolerance (value of tolerance on both the positive and negative sides of the ideal curve, within which 90% of the data points of real mechanism lie), t-test (to check if the two means are reliably different from each other), F-test and KolmogorovSmirnov test. All these six measures were computed for a set of preliminary simulations with varying levels of clearance. The mean deviation was found to be the most sensitive of these measures. Thus, for the rest of the analysis, we used mean deviation as a measure of deviation of the performance of the mechanism from the ideal mechanism." + ] + }, + { + "image_filename": "designv11_80_0003733_s00158-020-02741-x-Figure5-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003733_s00158-020-02741-x-Figure5-1.png", + "caption": "Fig. 5 The transition zone (third power function)", + "texts": [ + "8 m2, and m +M= 70 kg, then the differential equation is dv d\u03b8 \u00bc 1059:968cos\u03b8\u221252:998sin\u03b8\u22120:0508v2 v : \u00f011\u00de The relationship between the athlete\u2019s velocity and the incl inat ion of the arc was obta ined using Mathematica. v2 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u22121635:35e\u22120:1016 90\u2218\u2212\u03c6\u00f0 \u00de p \u00fe 319sin\u03c6\u00fe 2087:2cos\u03c6; where v2 is the skiing velocity at the end of the transition zone. The following relationship was present between skiing distance and inclination: s2 \u00bc R \u03c0 90\u2218\u2212\u03c6\u00f0 \u00de 180 \u00f012\u00de The inclination change of the arc corresponds to the change in skiing distance. s2 is the length of the transition zone (Arc) (Fig. 5). 2. Third power function In order to simplify the calculation, the start point was set as the origin of the coordinate axis. A local coordinate system was set up (\u03b7-\u03b6). The equation for the third power function is (Gasser 2008) \u03b7 \u00bc c\u03b63; \u00f013\u00de d \u00bc 2r1sin \u03b3\u2212\u03b1\u00f0 \u00decos2 \u03b3\u2212\u03b1\u00f0 \u00de: \u00f014\u00de l \u00bc \u222bd0 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 1\u00fe 9c2\u03b64 q d\u03b6 \u00bc d 1\u00fe 0:1tan2 \u03b3\u2212\u03b1\u00f0 \u00de ; \u00f015\u00de c \u00bc tan \u03b3\u2212\u03b1\u00f0 \u00de=3d2 \u00f016\u00de where \u03b3 is the inclination of the start zone of the inrun (\u00b0), and \u03b1 is the inclination of the takeoff zone of the inrun (\u00b0), r1 is the radius of E2 (m)" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001649_ccta.2019.8920549-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001649_ccta.2019.8920549-Figure1-1.png", + "caption": "Fig. 1: Current vectors in the rotating reference frame (dqframe) fixed to the two-pole rotor (blue) and the stationary reference frame (ab-frame) fixed to the stator (brown).", + "texts": [], + "surrounding_texts": [ + "Due to the variable reluctance of the motor, the inductance of a salient PMSM is dependent on the rotor position. The electrical dynamics of a salient two-phase PMSM model in the stationary reference frame is presented in [6], [7], [14]. The mechanical dynamics of the model remain the same as in the nonsalient model, hence (1c) and (1d) are still valid. L0 is the nominal inductance and \u2206L defines the amount of change of inductance due to saliency. L and \u2206L are given by L0 = Ld + Lq 2 , \u2206L = Ld \u2212 Lq 2 , (7) where Ld is the inductance in the d-direction and Lq is the inductance in the q-direction. The electrical dynamics of the salient two-phase PMSM in the rotating reference frame are presented in (8) Ld did dt Lq diq dt = [ vd vq ] + [ \u2212R N\u03c9mLq \u2212N\u03c9mLd \u2212R ] [ id iq ] + [ 0 \u2212Km\u03c9m ] . (8) As a consequence, the torque is now defined as (9) (see [19]). \u03c4em = Kmiq + (Ld \u2212 Lq)idiq (9)" + ] + }, + { + "image_filename": "designv11_80_0000647_s1028335819060041-Figure2-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000647_s1028335819060041-Figure2-1.png", + "caption": "Fig. 2. Frequency of nonlinear oscillations (9) and expression (11) as functions of the amplitude of oscillations . 1 is", + "texts": [ + " The substitution of this function into (7) makes it possible to express the function H0 from (4) through the variable : . The frequency of nonlinear oscillations is found from Eqs. (7) and (8): (9) The oscillations of the unperturbed system are given by (10) In the cases of oscillations in the neighborhood of the equilibrium state ( ) and separatrices (0 < ), frequency (9) takes values near zero. In these cases we have, respectively, For function has a maximum \u03bb* = 0.7786 which is achieved at q0 = = 2.0342 = 116\u00b030\u2032. The graph of function is shown in Fig. 2. Point is a unique point where the unperturbed system degenerates in the domain of oscillations; at this point the following value vanishes: (11) We present the graph of function (11) in Fig. 2. For value (11) is positive, and for it is negative. Domain of Rotations ( ) Consider rotations for p > 0 (case p < 0 is considered similarly). We introduce the variables by the relations (12) k \u2202 \u03c0= \u2202 \u22122 2 . 32 2 1 ( ) k I k k K k I =0 ( )H h I \u03bb \u2212 \u03c0\u2202 \u2202 \u2202\u03bb = = = \u2202 \u2202 \u2202 22(2 1) . 2 ( ) kh h k I k I K k = = \u03bb\u03be +0 0 0 0 , ( and are constats). I I w w I w < 00 1q \u03c0 \u2212 0 1q \u03c0 \u03c0\u03bb \u2245 \u03bb \u2245 \u2212 \u03c0 \u22120 0 2 2and . 2 ln( )4 ( 2 / 2) q qK < < \u03c000 q \u03bb 0( )q 0*q \u03bb =0 0*q q \u2202 \u2202\u03bb= \u2202\u2202 \u2212 \u2212 \u2212 \u2212= \u03c0 \u2212 \u2212 2 0 2 2 2 2 2 2 2 2 3 (1 )(4 1) ( ) (2 1) ( ) ", + " Domain of Oscillations ( ) We can show that Fourier coefficient (19) may be nonzero only for even m. With extremely laborious computations, we have found an explicit expression for it (23) where k is determined by Eq. (7) and d is determined by the relation = 1. From Eq. (23) it is clear that the sign of coefficient coincides with the sign of value . The computations show that in the interval the inequality is satisfied; therefore, the number of roots of the equation in this interval is equal to s \u2013 1. For small q0 and for q0 close to , we have, respectively, We take into account that (see Fig. 2) for small the value is positive and for close to it is negative and obtain that for small the studied subharmonic oscillations are stable for odd s and unstable for even s. If the values of are close to , then there is instability for any s. (We recall that we consider the number r in Eq. (22) even; for odd r the judgments about the stability will be the opposite). Figure 3 shows curves (20) for . The segments of the curves that correspond to the stable and unstable pendulum oscillations -periodic in are denoted by the solid and dashed lines, respectively" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000957_icma.2019.8816405-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000957_icma.2019.8816405-Figure1-1.png", + "caption": "Fig. 1. The overview structure of the Spherical robot", + "texts": [], + "surrounding_texts": [ + "A. Virtual prototype model establishment Model our spherical robots in the SolidWorks environment. Each individual component is designed and finally assembled into an overall model. After checking the model with SolidWorks, saved in Parasolid(*.x_t) format. Open Adams, enter into the GUI and click [file/import] in the menu bar, then select File Type as Parasolid (*.xmt_txt, *.x_t.....) to import the model. The second step sets the model properties. The parts of the model in SolidWorks was represented in the Adams model, but they can\u2019t move and it looks bare. We need to change the quality, torque, color and name of each part to meet our requirements. Because there is no actual servo motor in the simulation experiment, we need to add the rotating joint to the connections of the robot, and use the rotating joint motion to move in a specified function. In addition, we also need to add ground and give contact to the part to realistically reproduce the experimental scene. The ground plane was built to improve the simulation\u2019s third dimension. We can give the rotation a simple function to see if the robot can move." + ] + }, + { + "image_filename": "designv11_80_0001811_j.procir.2019.09.016-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001811_j.procir.2019.09.016-Figure3-1.png", + "caption": "Fig. 3. Parameters for a bended tape", + "texts": [ + " The aim of the algorithm is to determine the geometric parameters with which the robot should bend the tapes. The tapes must be bent in such a way that they are preformed as close as possible to the target shape. The input of the optimization is a CAD file of the surface of the target shape. For the described process the CAD file needs to be exported as STL file. The output of the optimization is a set of bending parameters with which the bending robot is able to bend the tapes. First, it is necessary to define the bending parameters. Therefore a syntax for the description of bended tapes is needed Fig. 3 shows an example of a tape which is bent two times. The bending parameters are determined with the trajectory of the tape\u2019s centre line. Firstly, \ud835\udc59\ud835\udc59\ud835\udc56\ud835\udc56 determine the lengths of the sections between the bending edges. The angles \ud835\udefc\ud835\udefc\ud835\udc56\ud835\udc56 describe the rotation of the bending edge, since oblique-angled bending is considered. Finally, the angles \ud835\udefd\ud835\udefd\ud835\udc56\ud835\udc56 describe how far the tape is bent around the bending edge. Therefore, a tape with \ud835\udc5b\ud835\udc5b bends can be described with (3\ud835\udc5b\ud835\udc5b + 1) bending parameters. In order to optimize the bended tape, its position and geometry need to be compared to the target shape" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002057_s00006-019-1039-z-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002057_s00006-019-1039-z-Figure3-1.png", + "caption": "Figure 3. a The Delta manipulator. b A network representation of the Delta manipulator", + "texts": [ + " Figure 2a, b below shows an example of a typical 3 limbed manipulator and its position vectors \u2018iXb n\u2019 and \u2018iXg n\u2019 for i = (1, 2). The i-th kinematic pair\u2019s screw (generally given with the symbol $i [35]) which will be denoted by \u2032X \u03b8 i n+1 \u2032 is given as X \u03b8 i n+1 = [Xi n+1 \u00d7 si, si ] (2b) Where si is the i-th kinematic pair\u2019s joint axis and Xi n+1 is the linear position vector from joint \u2018i\u2019 to the end-effector. We demonstrate two examples of the computation of X \u03b8 i n+1 , with a Delta manipulator presented in Fig. 3. It has the following kinematic pairs (in meters), where Hi is the initial vector of the kinematic pair from joint \u2018i\u2019 to \u2018i+1\u2019, and \u2018si\u2019 is the initial unit vector along the joint axis: H0 = {0.151, 0, 0} H1 = {1.5, 0, 0} , H2 = {0, 0, 0} , H3 = {2.5, 0, 0} , H4 = {0, 0, 0} , H5 = {0.25, 0, 0} H6 = {1.5, 0, 0} , H7 = {0, 0, 0} , H8 = {2.5, 0, 0} , H9 = {0, 0, 0} , H10 = {0.25, 0, 0} H11 = {1.5, 0, 0} , H12 = {0, 0, 0} , H13 = {2.5, 0, 0} , H14 = {0, 0, 0} , H15 = {0.25, 0, 0} s1 = [0, 1, 0] , s2 = [0, 0, 1] , s3 = [0, 1, 0] , s4 = [0, 0, 1] , s5 = [0, 1, 0], s6 = [\u22120" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001548_1.5132177-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001548_1.5132177-Figure1-1.png", + "caption": "FIGURE 1. The general scheme of layer-by-layer growth of a sample by wire-feed electron beam additive manufacturing (a) and the trajectory of motion for the construction of each layer (b): 1 - movable table; 2 - titanium substrate; 3- electron gun; 4 - electron beam; 5 - wire feeder; 6 - wire; 7 \u2013 weld pool; 8 - grown material", + "texts": [ + " As initial material the wire from Ti-6Al-4V titanium alloy with diameter of 0.8 mm was used. The samples were grown on a technical titanium substrate of 75 \u00d7 75 mm thickness of 2.5 mm mounted on a mobile worktable capable of moving in 3 directions along the X, Y and Z axes. During the growth process, the melt pool was formed by scanning an electron beam in the form of a ring with a diameter of 5 mm. The scanning frequency was 1 kHz. The wire was fed by a special feeder fixedly mounted relative to the electron gun (Fig. 1a). The combination of the parameters of the electron beam, wire feed rate and Proceedings of the International Conference on Advanced Materials with Hierarchical Structure for New Technologies and Reliable Structures 2019 AIP Conf. Proc. 2167, 020310-1\u2013020310-5; https://doi.org/10.1063/1.5132177 Published by AIP Publishing. 978-0-7354-1912-4/$30.00 020310-1 linear growth rate (the rate of movement of the table relative to the beam) were chosen in such a way as to ensure the stability of the process of melting the tip of the wire and draining the liquid metal into the molten bath in a continuous flow. When growing the sample, 3 layers were successively deposited on the substrate without mutual displacement in the XY plane according to the scanning strategy shown in Figure 1, b. Thus, a sample of Ti-6Al-4V alloy with a size of 30 \u00d7 30 \u00d7 2.4 mm was obtained, which consisted of three identical layers. For metallographic analysis, the samples were ground and polished according to the standard procedure, then etching in a reagent with the composition 2% HF, 2% HNO3, 96% H2O (volume fractions indicated) was carried out. The microstructure was investigated using optical (OM) and scanning electron microscopy (SEM). The phases were identified on a DRON-7 X-ray diffractometer using CoK\u03b1 radiation with a step of 0" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0002245_j.automatica.2020.108895-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0002245_j.automatica.2020.108895-Figure1-1.png", + "caption": "Fig. 1. The geometric relationships between angles of bearing vectors and a subtended angle.", + "texts": [ + " Find the bearing vector gvu, for each edge (u, v) \u2208 E , using communications in G and the measurement set S. Consider the setting of Problem 1 with a network of N agents and their connected communication graph G = (V, E). Let \u0338 gvu \u2208 [\u2212\u03c0, \u03c0) be an angle of the bearing vector gvu with respect to the global coordinate frame g \u2211 . Then, for each edge (u, v) \u2208 E , estimating the angle \u0338 gvu is identical to finding the bearing vector gvu in Problem 1, since gvu = [cos (\u0338 gvu) , sin (\u0338 gvu)] T . The angle \u0338 gvu can be estimated by the geometric relationships between itself and \u0338 gwv , shown in Fig. 1, as follows: \u0338 gvu = \u23a7\u23a8\u23a9 \u0338 gwv \u2212 ( \u03b1v wu + \u03c0 ) if w \u0338= u and \u03b1v wu \u2208 S, (a) \u0338 gwv \u2212 \u03c0 if w = u. (b) (1) Our main approach to estimate \u0338 gvu for each (u, v) \u2208 E is to use the consensus protocol based on the geometric relationships in (1), as follows: First, using the concept of directed line graph in Definition 2, each edge (u, v) \u2208 E will be mapped to a virtual agent (u, v)\u03f5 , called edge agent, whose orientation angle is identical to \u0338 gvu. Then each edge agent (u, v)\u03f5 will receive the angle information from its neighbors and estimate its orientation angle through consensus protocol" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0003864_ies50839.2020.9231879-Figure6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003864_ies50839.2020.9231879-Figure6-1.png", + "caption": "Figure 6. Testing the default position on the uneven surface", + "texts": [ + " In this system there are two sensors, the first sensor is a load cell sensor to determine the position of the pressure center at the robot's foot. When the robot is on an uneven surface, the robot will be able to know the level of flatness of the pedestal base based on the pressure center. From this pressure center value will be used as a correction from the robot's default position. While the second sensor is a gyroscope sensor that functions as an angular acceleration sensor, which is used as the center of mass o f the LIPM control. I II . E x p e r im e n t s In this test, we will divide it into two tests. The first test is shown in Figure 6. The first test is testing the pose or default condition of the robot, to determine the robot's ability to maintain balance when on an uneven surface. The robot will be placed on an adjustable board slope, so that the robot will know the response to make corrections to the uneven surface when the default position conditions. 215 Authorized licensed use limited to: University College London. Downloaded on November 02,2020 at 04:47:10 UTC from IEEE Xplore. Restrictions apply. different from previous tests" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure90.6-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure90.6-1.png", + "caption": "Fig. 90.6 Distribution of Von Mises stress a and plastic strain b at plasticity stage in liner", + "texts": [ + " The stable bond between liner and composite overwrap shall fail in case of design irregularities, complex curvatures in liner, thus creating a debond region. In the present analysis, a debond region is created by removing the contact interaction between liner and composite overwrap to study the ill-effects on liner. The eigenvalue linear analysis is performed for determining the load at which liner buckles during depressurization with subsequent buckling modes. The burst load of 4.32 MPa applied on inner surface of liner at plastic stage showed severe yielding with Von Mises stress of 309 MPa (refer Fig. 90.6a) and plastic strain of 0.76% (refer Fig. 90.6b) at the transition regions near to pole and equator region. Further, the first failure location of liner is seen at transition region near the 1080 R. Pramod et al. equator section. Failure is observed in terms of crack appearing in liner and complete rupture is observed as equator region gets completely detached from the liner section at Von Mises stress of 312 MPa (refer Fig. 90.7a) and plastic strain of 0.82% (refer Fig. 90.7b). From the analysis, it was also required to know, if the failure is occurring in weld zones of liner as weld specimen has lower strength (refer Table 90" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0000000_cac.2018.8623073-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0000000_cac.2018.8623073-Figure3-1.png", + "caption": "Fig. 3. Depiction of the multibody representation", + "texts": [ + " (ii)The aerodynamic forces and the aerodynamic torque are changed with aerodynamic parameters\u2019 great change. (iii)Bring Serious inertial forces and torque disturbances. Hence, comparing with conventional aircraft, the morphing will bring new features to the modeling and control system, which needs further researches. The definition of wingspan and sweep rate are / 45\u03bb = \u039b \u00b0 and 1/ l\u03be = \u0394 , hence [0,1]\u03bb \u2208 and [0,0.8]\u03be \u2208 , in this equation, \u039b is the wing sweep angle, \u0394 is the length of telescopic outer wingspan. The values of other variables tagged in Fig. 3 are shown in Table I. This paper only consider the longitudinal motion of morphing aircraft. The flight condition is selected as the altitude 9144h m= and the Mach number a 0.5M = ( 0 170.15m/sV \u2248 ). A. Nonlinear Modeling by Kane Regardless of the variant aircraft, the aerodynamic shape needs to be changed within a fairly large range. Therefore, the single-rigid assumption assumed by conventional aircraft modeling can\u2019t be satisfied. Therefore a multibody modeling method should be introduced in the modeling" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001314_978-981-32-9072-3_65-Figure82.3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001314_978-981-32-9072-3_65-Figure82.3-1.png", + "caption": "Fig. 82.3 2D model drawing of Model 2 (push fit type with cross-drills)", + "texts": [], + "surrounding_texts": [ + "See Figs. 82.1, 82.2 and 82.3 and Tables 81.1 and 82.2. 82.3.3 Software Used 1. Modelling\u2014SOLIDWORKS 2016 2. Analysis\u2014ANSYS WORKBENCH R16. 82 Finite Element Analysis of a Disc Brake Mounted on the Axle \u2026 975 976 E. Madhusudhan Raju et al." + ] + }, + { + "image_filename": "designv11_80_0003328_012034-Figure3-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0003328_012034-Figure3-1.png", + "caption": "Figure 3. Structure components.", + "texts": [ + " Therefore, a failure theory, involving the factor of safety and the maximum stress, is applied for the determination of the Sy. It is calculated with the Equation (4). \u03c49\": = ;$ (. . (4) Where, \u03c49\": is the maximum shear, Sy is the yield strength, n is the safety factor. The parts are sized according to the conditions of the machine. The vertical structure of 250 mm, with a space for the welding gun that rests on a base connected to the pair of shafts, and the location of the motor that will give the torque to the power screw. Figure 3 shows a preview of the design using SolidWorks software. 6th International Week of Science, Technology and Innovation (6th IWSTI) Journal of Physics: Conference Series 1587 (2020) 012034 IOP Publishing doi:10.1088/1742-6596/1587/1/012034 In the application of the stress equations, the selected values, and the use of the commercial tables of the materials, allowed finding and calculating the following factors exposed in Table 3. In accordance with the results of subsection 3.1, the maximum stress on the power screw is 283" + ], + "surrounding_texts": [] + }, + { + "image_filename": "designv11_80_0001981_s11148-020-00406-2-Figure1-1.png", + "original_path": "designv11-80/openalex_figure/designv11_80_0001981_s11148-020-00406-2-Figure1-1.png", + "caption": "Fig. 1. Diagram of a dragline excavator (a) during its movement and a cross-section of the trailing edge niche of the base with shoes (b).", + "texts": [ + " To implement a dragline step, it is necessary to raise the front edge of the base in the direction of travel with lifting hydraulic cylinders located on the shoes. Traction hydraulic cylinders, on the other hand, operate to ensure the direct movement of the dragline. Such maneuvers are very energy-intensive; therefore, a model was proposed in [2], which reduces the cost of moving the dragline due to a constructive change in the mechanism and the use of hydraulic accumulators. The proposed device for moving walking machines (see Fig. 1) contains a rotary housing 1 with shoes 2. The ground 3 support base 4 has an edge 5. A shoe 7 is built into a niche 6. The support base 4 is made with the possibility of horizon- tal rotation around its axis 8 when the case 1 raised on the shoes 2 is stationary. To raise the rear of the edge 5 of the support base 4 above the ground 3, the niche 6 is made 558 1083-4877 20 06006-0558 \u00a9 2020 Springer Science+Business Media, LLC 1 Ural State Mining University, Yekaterinburg, Russia. 2 Magnitogorsk State Technical University named after G" + ], + "surrounding_texts": [] + } +] \ No newline at end of file