[ { "image_filename": "designv11_22_0003158_9781315768229-Figure9.20-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure9.20-1.png", "caption": "Figure 9.20 Two-point travelling steady", "texts": [ " To overcome this, live or rotating tailstock centres are available, the centres of which run in bearings which will withstand high pressures without overheating. Tailstock centres are often required for long work which is held in the chuck but requires support owing to its length. If unsupported, long slender work may tend to be pushed aside by the forces of cutting. To overcome this, a two-point travelling steady is used which provides support to the workpiece opposite the tool as cutting is carried out along the length of the work, Fig.\u00a09.20. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 9 137 9 Turning does not have to be removed to load and unload workpieces. Accurate turning of plain diameters and faces can be simply carried out on a centre lathe. Wherever possible, diameters should be turned using the carriage movement, as the straightness of the bed guideways ensures parallelism of the workpiece and power feed can be used. Avoid using the top slide for parallel diameters, since it is adjustable for angle and difficult to replace exactly on zero without the use of a dial indicator" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003373_5.0035987-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003373_5.0035987-Figure1-1.png", "caption": "FIGURE 1. Schematic diagram of the rocket engine combustion chamber 1 \u2013 case; 2 \u2013 nozzle head; 3 \u2013 cylindrical section; 4 \u2013 nozzle;", "texts": [ " The combustion chamber is designed to convert combustible and oxidizing generator gas, as a result of chemical reactions, into high-temperature combustion products that generate reactive force - engine thrust at outlet from the nozzle. The combustion chamber is a brazed-welded construction consisting of a gas duct with a gas distribution grid, a flat mixing head, a block of the combustion chamber, a profiled upper nozzle block and a lower nozzle block. The combustion chamber block is an integral soldered construction consisting of two parts: the inner wall (shell) and the outer wall (shirt). Figure 1 shows a schematic diagram of the rocket engine combustion chamber. AIP Conf. Proc. 2318, 030004-1\u2013030004-7; https://doi.org/10.1063/5.0035987 Published by AIP Publishing. 978-0-7354-4060-9/$30.00 030004-1 In the manufacture of the combustion chamber shell, it is required to ensure high accuracy and quality of the heatloaded part surface layer. Analysis of defects and failures of aircraft and space technology products showed that the main reasons for the violation of these factors are design and technological flaws, as well as basic material supplied by third-party organizations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000680_iros40897.2019.8968461-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000680_iros40897.2019.8968461-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_22_0003676_s00170-021-07413-8-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003676_s00170-021-07413-8-Figure6-1.png", "caption": "Fig. 6 Components of the die set", "texts": [ " On the other hand, in the free area of the side wall region, it is difficult to prevent that phenomenon. In this paper, to investigate the factors and parameters affecting wrinkling and make wrinkling easier to observe, the shape of the punch is designed to be a cylinder rather than a cone. This technique was used by other research works, such as [11], for the deep drawing process. It should be stated that in this way, the final part is formed as a cone which also serves the secondary purpose of forming a conical workpiece. Figure 6 shows the components of the die and the assembled set on the testing machine. A variable hydraulic pump with the maximum pressure of 50MPa and a constant flow pumpwith the maximum pressure of 100 MPa were used to supply the fluid in the die cavity and blank rim areas, respectively. The SAE10 hydraulic oil with a viscosity of 5.6 CST was used as the pressure medium that provides the required pressure for the forming process. Selecting a fluid with suitable viscosity helps to create the radial pressure [33]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002384_s12206-020-0325-y-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002384_s12206-020-0325-y-Figure8-1.png", "caption": "Fig. 8. Bulgy and hollow shapes (steel side): (a) Initial model; (b) optimum model.", "texts": [ " 8 and 9 that the volume of aramid/phenol is enlarged after optimization as the steel volume decreases. Steel and composite parts are jointed together as in Fig. 9. Therefore, the decreased steel volume means the increase in a composite volume. Maximum von Mises stresses are summarized in Table 4 for both steel and composite sides before and after optimization. Due to the increase in composite volume, the peak stress is reduced below the phenol strength, 100 MPa. In addition, the maximum stress in a steel surface is also decreased, due to the fillet radius occurred in the middle height, as in Fig. 8. Two kinds of helical gears, an aramid/phenol and 100 % steel gears, were fabricated for vibration tests to investigate the damping effect of the composite part inserted between inner and outer steel parts. Force is applied to a tooth of two helical gears in a tangential direction by using a shaker, as shown in Fig. 11. The experimental setup for vibration test is depicted in Fig. 11. Force is imparted to the point A on a tooth surface and then acceleration is measured at the point B, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002736_jsen.2020.3022421-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002736_jsen.2020.3022421-Figure6-1.png", "caption": "Fig. 6 (a) Bending deformation model of the soft finger without silicone joints based on ABAQUS; (b) Bending deformation model of the soft finger with silicone joints based on ABAQUS", "texts": [ " Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. designed soft finger model, the 3D model of soft finger in SolidWorks was imported into the FEM software ABAQUS. Then we set the material properties of each component through the parameters in Table \u2160. We also completed the assembly of the soft finger model, and arranged the model grid, added the driving torque and boundary fixation constraints. Fig. 6 is the 3D model of soft finger simulation. We conducted the deformation experiment of the soft finger without silicone joints and the finger with silicone joints in Section \u2163. In the experiment, the soft finger with three optical fibers was fixed on the aluminum alloy frame, and the SMA-3 wire was connected to the tensile equipment with the pressure sensor. The driving force of the soft finger was changed through the rocker, while the changing curvature was measured. Then the working performance of the FBG sensor was analyzed through our theoretical research" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure12-1.png", "caption": "Fig. 12. Stress analysis results of the pinion with single contact point (unit: MPa): (a) von Mises stress; (b) shear stress.", "texts": [ " 11, the maximum contact stress between the internal gear pair is 1429.4 MPa. The maximum stress occurs at contact point, which is located on the middle of the tooth profile. It has regular elliptical distribution along the direction of tooth width, and the distribution area has the trend of expanding to the tooth root direction. With the increase of the contact area, the contact stress will gradually decrease. The maximum von Mises stress and shear stress of the pinion with single contact point in Fig. 12 are 729.02 MPa and 374.89 MPa, respectively. The maximum von Mises stress and shear stress of the internal gear with single contact point in Fig. 13 are 729.58 MPa and 201.07 MPa, respectively. 4.2 Stress analysis with two contact points Similarly, we established a gear model with two contact points according to the proposed generation methods of tooth profiles. It means that there are two contact points on one contact pair at the same time during the meshing process. Here, the meshing pair should also be the pinion with convex tooth profile and the internal gear with concave tooth profile" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001598_iecon.2015.7392524-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001598_iecon.2015.7392524-Figure8-1.png", "caption": "Fig. 8: Division of the Preisach plane into N D 12 equally-distanced values for \u02db and \u02c7 , and the resulting 12 discrete levels obtained on the hysteresis curve.", "texts": [ " By considering an N number of memorized extrema, the piezoelectric expansion becomes: x.t/D 8\u0302 \u02c6\u0302\u0302\u0302 \u02c6\u0302< \u02c6\u0302\u0302\u0302 \u02c6\u0302\u0302: X.u.t/;\u02c7N /C C NP iD1 .X.\u02dbi ;\u02c7i 1/ X.\u02dbi ;\u02c7i // Pu.t/ > 0I X.\u02dbN ;\u02c7N 1/ X.\u02dbN ;u.t//C C NP iD1 .X.\u02dbi ;\u02c7i 1/ X.\u02dbi ;\u02c7i // Pu.t/ < 0: (10) In order to obtain the values needed for each Everett function, and consequently the discrete Preisach model, the previouslydefined limiting triangle T0 is now split up into a mesh containing values for the point pairs f.\u02dbi ;\u02c7j /ji D 1::N;j D i::N g. Fig. 8 shows the geometrical interpretation of the discrete Preisach plane on the left hand side, together with the obtained input-output representation on the right-hand side, for an arbitrary value of N D 12. The number of \u02db and \u02c7 values equals the number of discrete steps on the hysteresis curve. Naturally, the higher N is chosen, the finer the mesh becomes and consequently, the finer the hysteresis representation will become. This is at the expense of exponentially increasing memory requirements to store all the values" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000412_012017-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000412_012017-Figure5-1.png", "caption": "Figure 5. Determination of the boundary condition and loading", "texts": [ " The simulation of collected data was done on the bolt hole area on leaf springs number eight. The usage of hole area due to the fracture failure is around it. From the result of simulation, the stress distribution known, as well as stress intensity factor and strain also obtained [14,15]. The steps and simulation process for numerical analysis were done using FEM software by giving fixed support on the spring eye. The loading type used was centralized load given on the bottom of the spring. The truck loading can be calculated by analytical analysis as shown in Fig. 5. The meshing steps used the obtained nodes number of 633438 and the number of elements of 5280. The meshing result of finite element analysis for all spring is shown Fig. 6. IC-DAEM 2018 IOP Conf. Series: Materials Science and Engineering 547 (2019) 012017 IOP Publishing doi:10.1088/1757-899X/547/1/012017 The results of hardness test areshown in Fig. 7 horizontal axis from point A to B and vertical axis from C to D in the sum of 22 points. The distance between point approximately 2 to 5 mm (ref" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002758_j.jbiomech.2020.110050-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002758_j.jbiomech.2020.110050-Figure2-1.png", "caption": "Fig. 2. Anatomical coordinate system assigned to the residual femur bone using landmark locations. Coordinate systems for the residual femur were established from the three-dimensional models generated after segmentation as follows: Of: Most distal point of the residual femur during CT scan. Yf: Vector connecting the distal end of the residual femur to the greater trochanter. Zf: Vector perpendicular to plane formed by Yf and the center of the femoral head pointing anteriorly for the left side and posteriorly for the right side. Xf: Vector perpendicular to both Yf and Zf.", "texts": [ "7 and shown to correlate with Short-Form 36 (SF-36)-Item Health Survey for validity of the measurements (Hagberg et al., 2004). A previously validated volumetric model-based tracking technique, with in vivo accuracy of 0.7 mm or better in translation and 0.9 or better in rotation, was used to match digitally reconstructed radiographs created from the subject-specific CT scans to the biplane radiographs (Fig. 1e) (Anderst et al., 2009). Coordinate systems for the femur were established using anatomical landmarks (Fig. 2) and the socket coordinate system was established using surface marker data from the static standing trial (Fig. 3). A separate local coordinate system was established to track the socket during the walking trials using four markers placed on the socket and the medial/lateral knee markers. Up to four variations of these local coordinate systems were created, each using one of the four socket markers and the two knee markers (Fig. 3) to track the socket even if a marker was not visible during a walking trial" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002396_s12206-020-0315-0-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002396_s12206-020-0315-0-Figure4-1.png", "caption": "Fig. 4. Geometric model and mesh generation.", "texts": [ " The change trend is generally non-linear. The expression can be written as follows: [ ( )]{ } [ ( )]{ } { ( )}c T T K T T q T+ =& (8) K(T) Is the conduction matrix (heat coefficient, convection coefficient, radiation and other parameters) changing with temperature; c(T) Is the specific heat matrix with temperature change, which is related to the internal energy of the system; {q(T)} Is the node heat flux vector with temperature change. The laser cladding geometry model and mesh division are shown in Fig. 4. The matrix size is 40\u00d730\u00d78 mm. In order to improve the operation speed, symmetrical model is adopted, and the details such as rounded corners of parts are neglected. In order to improve the accuracy of calculation, the meshes near the cladding layer are dense, and the mesh size of the matrix far from the cladding layer increases gradually. And the matrix is divided into free meshes. By applying temperature load to the element, the transient analysis with small time step is carried out. The analysis results are taken as the initial conditions of temperature" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003845_s00521-021-06342-7-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003845_s00521-021-06342-7-Figure1-1.png", "caption": "Fig. 1 The sample model of a NWMR", "texts": [ " And 9pm; qm; pM; qM [ 0 such that pmIn P\u00f0t; x\u00de pMIn; qmIn Q\u00f0t; x\u00de qMIn Lemma 1 [37] Under Assumptions 1 and 2 where r[ 0 is an arbitrary fixed constant, system (4) is ULES for 8\u00f0t0; xi0\u00de 2 R\u00fe Br, if there exists positive constants a and T0 such that Z t\u00feT0 t \u00bdB\u00f0s; x\u00f0s; t0; x0\u00de\u00deB\u00f0s; x\u00f0s; t0; x0\u00de\u00deT \u00fe D\u00f0s; \u00f0s; t0; x0\u00de\u00de ds aIm; 8t t0 \u00f05\u00de Lemma 2 [27] Consider a time-varying bounded block diagonal matrix B\u00f0t; v\u00f0t\u00de\u00de : \u00bdt0;1\u00de RLl ! RLm Ln where Bi\u00f0t; vi\u00f0t\u00de\u00de : \u00bdt0;1\u00de Rm n\u00f0i 1; :::L\u00de and the Laplacian matrix L 2 RL L of an undirected connected graph, assuming that Bi\u00f0t; vi\u00f0t\u00de\u00de is cooperative-ul PE, then Z t\u00feT0 t \u00bdB\u00f0s; v\u00f0s\u00de\u00deB\u00f0s; v\u00f0s\u00de\u00deT \u00fe q\u00f0L Im\u00de ds aILm \u00f06\u00de where q is a positive constant. Considering a multi-robot system composed of L \u00f0L[ 1\u00de nonholonomic wheeled mobile robots with the same mechanical and electrical structure (see Fig. 1), each of which can be modeled as [38] _qi \u00bc S\u00f0qi\u00degi \u00f07\u00de M\u00f0qi\u00de _gi \u00fe C\u00f0qi; _qi\u00degi \u00fe X\u00f0qi\u00degi \u00fe G\u00f0qi\u00de \u00fe F\u00f0qi\u00de \u00fe sdi \u00bc B\u00f0qi\u00deui \u00f08\u00de where i 2 I\u00bd1; L , qi \u00bc \u00bdxci; yci; hi T and gi \u00bc \u00bdvi;xi T represent the pose and velocity vector of the ith mobile robot system, respectively. S\u00f0qi\u00de 2 R3 2 represents the kinematic matrix, M\u00f0qi\u00de 2 R2 2 denotes the bounded positive definite symmetric matrix, C\u00f0qi; _qi\u00de 2 R2 2 is the vector of Coriolis and centripetal forces, X\u00f0qi\u00de 2 R2 2 is the velocity transformation matrix, G\u00f0qi\u00de 2 R2 1 denotes the vector of gravitational torque, F\u00f0qi\u00de 2 R2 1 is the friction vector, sdi 2 R2 1 is the bounded external disturbance, B\u00f0qi\u00de 2 R2 1 represents the input transformation matrix and ui \u00bc \u00bdui1; ui2 T is the voltage input" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000514_j.mechmachtheory.2019.103667-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000514_j.mechmachtheory.2019.103667-Figure9-1.png", "caption": "Fig. 9. Field of equivalent von Mises stresses on the face-gear tooth surfaces (a) without considering supporting shafts and (b) with supporting shafts being considered.", "texts": [ " When shafts are not considered, areas of high contact stresses appear on the pinion and face-gear tooth surfaces. Those areas are located in the upper corner of the inner section of the active tooth surface of the face gear and the corresponding area in the pinion tooth surface. However, when shafts are considered into the finite element model, the contact pattern is shifted towards the outer radius of the active tooth surfaces of the face gear and those areas of high contact stresses appearing in the inner radius are avoided and the maximum contact stress reduced. Fig. 9 shows the field of von Mises stress at contact position eleven for the case when supporting shafts are not considered or, in other words, for ideal conditions of meshing (Fig. 9(a)) and for the case when elastic shafts are considered (Fig. 9(b)). Results shown in Fig. 9 suggest that the active tooth surfaces of the face gear should be trimmed at the inner and outer upper corners of the surface to avoid those areas of high contact stresses. The discrepancy between the maximum stress shown in Fig. 9 and the stress represented in Fig. 8 is due to fact that the field of von Mises stresses is based on the averaged value of nodal stresses, where as the stresses shown in Fig. 8 (and following ones directed to show the variation of stresses along the 21 contact positions) are based on the direct reading of the stress at the integration points of the finite elements. Fig. 10 (a) shows the theoretical boundaries of the active tooth surfaces of the face gear considered in the stress analysis shown above and Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002689_j.robot.2020.103609-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002689_j.robot.2020.103609-Figure1-1.png", "caption": "Fig. 1. An illustration of a string of three tethered drones and the ground station inertial reference coordinates.", "texts": [ " Section 4 includes the description of the ontrol scheme, followed by numerical simulations that are preented in Section 5. Some fundamental practical aspects are disussed in Section 6, and Section 7 concludes the paper. A stability roof and detailed model derivation are given in Appendices. . System description Consider a taut tether connected to an active ground station n its one end, and to a flying drone on its other end. Additionally, assume a group of N \u2212 1 drones, capable of moving freely along the tether (i.e., a total of N drones), as in Fig. 1. The ground station epresents the inertial reference frame. In the ground station, he tether is wrapped around a drum, driven by a DC motor. he tether length can be measured and adjusted by rotating he motor. Next, the tether passes through a 2-axis gimbal in he ground station. Each of the axes is connected to a rotation ngle sensor, which eventually provides a measure of the tether\u2019s ttitude with respect to the inertial frame [9]. Onboard each of the drones, the tether passes through a pulley nd a pair of in and out gimbals (see Fig", " (d) ether elongation is negligible, and (e), tether tension due to orsion is negligible. The dynamical model is developed with respect to the availble measurements, i.e., the tether-sections\u2019 length and attitude. or any number of N serially-connected tethered drones, the osition xn \u2208 R3 of the nth drone with n \u2208 [1,N] is given by, n = n\u2211 i=1 riqi (1) here, qi \u2208 S2 is a unit vector denoting the tether attitude, 2 = { q \u2208 R3: \u2225q\u2225 = 1 } and ri \u2208 R is a scalar denoting the tether ength along the unit vector qi (see Fig. 1). Using this, the instantaneous motion of drones can be govrned by two elements, one that is aligned with the tether beow and another that is perpendicular to it. The later element nd its influence on the position of the drone can be analyzed sing the theory of a rigid-body moving on a two-sphere S2, hile the aligned element changes the radius of the two-sphere. he angular state of the drone is expressed as a rotation marix R in the Special Orthogonal Group, defined by SO (3) \u225c R \u2208 R3\u00d73: RRT = I3\u00d73, det (R) = 1 } ; it represents the rotation of the body frame relative to the inertial frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000039_0954405419828590-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000039_0954405419828590-Figure4-1.png", "caption": "Figure 4. Schematic configuration of specimen orientations and dimensions of room temperature tensile tests: (a) position of tensile specimens, (b) dimension of tensile specimens in mm, (c) deposited wall and (d) tensile specimens in different areas.", "texts": [ " Moreover, microhardness testing was also performed parallel to the travel direction (Y) in longitudinal section (Y\u2013Z plane). The orientations and locations of the indentations are presented in Figure 3(b). All indentations in Figure 3 were measured with an interval of 0.5mm. Furthermore, the microhardness of each point was determined for three times. Tensile properties at room temperature of the deposited walls were measured in both the deposition height direction (Z) and travel direction (Y). Figure 4 gives the location, direction, section shape and geometrical size of the testing specimens. At first, the tensile properties between the travel direction (Y) and deposition height direction (Z) were compared. The locations of specimens are in the middle region of the wall, as shown in Figure 4(a). The tensile specimens in each orientation were prepared for three times. Second, to reveal the nonuniform tensile properties within the deposited part, tensile specimens were prepared in the travel direction (Y) from the near-substrate region to the top region with an interval about 10mm, as indicated in Figure 4(c) and (d). To reduce experimental errors, the tensile specimens in the same layer were prepared for three times in the travel direction (Y), and the tensile testing values were averaged. To study the mechanical properties as a function of location and orientation in detail, especially as shown in Figure 4(d), we must design a new and small dimension of tensile specimens due to a small width of the thin-walled part presented in Figure 2(b). Thus, the geometrical size of the tensile specimens was designed according to the standard of GB/T2282002 in China; however, the width was set at a relatively small value. Based on the testing standard of GB/T2282002, room temperature testing was conducted using an electronic universal testing machine (CMT4304) at an extension rate of 0.5mm/min. Figure 5 shows the typical macrostructure of a crosssection (X\u2013Z plane) in the deposited wall", " According to the plastic deformation theory, keeping the part volume constant, the number of grains, the volume fraction of grain boundaries and the dislocation obstacles decrease with the increasing grain size, resulting in the decreased metallic plastic deformation resistance and microhardness. In the top region, the microstructure consisting of granular bainite and acicular ferrite presents hardness without the heating effects from subsequent layers. Figure 12 shows the room temperature tensile properties of the walls in the deposition height direction (Z) and travel direction (Y). Three test specimens were applied for each direction, as displayed in Figure 4(a), and the error bar is given in Figure 11(a). Even three same positions for each direction are obtained by fabricating three walls deposited with same process parameters, and the tensile strength results are impossible to be the same due to inevitable measuring error from machining accuracy of specimens or electronic universal testing machine. It is shown that the ultimate tensile strength of deposited layers in the travel direction (Y) is slightly larger than that in the deposition height direction (Z)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure2.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure2.4-1.png", "caption": "Fig. 2.4 Visualization of the three important poses in the workspace of a bin-picking robot. The object pose WPO which defines the pose of the located object (yellow) in the world, the grasp pose GPO which described the pose of the grasped object (purple) in the gripper and the gripper pose WPG which is the pose of the gripper in the world (green) at which an object can successfully be grasped. It is assumed that the robot as well as the optical scanner are calibrated w.r.t. the world frame W . When the gripper grasps an object, the three transformations build a closed chain", "texts": [ " . 101 Own Publications and References . . . . . . . . . . . . . . . . . . . . . . . . . . . . 113 x Contents Figure 2.1 Results of Ikeuchi\u2019s and Horn\u2019s work on needle maps . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6 Figure 2.2 The used matched filter kernel of Dessimoz. Image taken with permission from [25] . . . . . . . . . . . . . . . 7 Figure 2.3 Results of Dessimoz\u2019 work on gripper pose estimation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 Figure 2.4 Visualization of the three important poses in the workspace of a bin-picking robot . . . . . . . . . . . . . . . 11 Figure 3.1 Regular structure of a 3D mesh acquired by a laser line scanner. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Figure 3.2 Rotation and translation invariant features of a dipole . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Figure 3.3 The assumption of a tangential contact between two oriented point pairs can be used to define a relative transformation ATB ", "2 Three different poses are important in a bin-picking system: \u2022 The object pose WPO , \u2022 the gripper pose WPG and \u2022 the grasp pose GPO . The object pose WPO hereby describes the pose of an object with respect to some world coordinate system.3 The gripper pose WPG is the goal pose of the end effector w.r.t. the world coordinate system to grasp an object. The grasp pose GPO is the pose of the object w.r.t. the gripper coordinate system, which is especially important for a defined placement of the object. These three poses build a transformation chain (see Fig. 2.4). Therefore, if two of them are estimated, the computation of the third is trivial. Beyond pose estimation, a second important problem arises. This problem can be named as \u201ccollision avoidance\u201d. Each gripper pose has to be analyzed for possible collisions of the gripper with the environment. Only when the pose estimation and collision avoidance is solved, can robot movements be executed safely. 2Within this work, it is assumed that only one type of objects is in the bin. Otherwise, an object detection step has to be executed additionally" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001509_j.procs.2015.12.319-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001509_j.procs.2015.12.319-Figure3-1.png", "caption": "Figure 3.Turning radius calculation model", "texts": [], "surrounding_texts": [ "For each of different types of conditions that will triggered the system to perform certain movement, poses or actions, there is a taken time for the system to return to its original condition. This means the system is finishing performing a certain loop and getting ready for another loop in the program flow. The time taken need to be recorded to know the response time of the system in order to know how well the develop platform react in an environment." ] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure8-1.png", "caption": "Fig. 8. Backlash caused by tooth thickness.", "texts": [ " Therefore, the simplified meshing stiffness of the helical gear can be expressed as ( ) ( )0=mk t k L t , (43) where k0 is the meshing stiffness density in a unit of contact line and L(t) is the total length of the contact line. Theoretically, the process of gear meshing should be a backlash-free meshing. However, in the actual manufacturing process, a backlash between the teeth profiles is necessary to prevent gear stuck caused by the error of gear manufacturing, machining, and installation, as well as the thermal expansion of the gear teeth. The size of the backlash can be changed by adjusting the length of common normal line or the center distance. As shown in Fig. 8, the backlash as a result of the lesser actual tooth thickness than the ideal non-backlash tooth thickness is given by = \u2212t i ab t t , (44) where ti is the tooth thickness of ideal non-backlash and ta is the actual tooth thickness. With the consideration of eccentricity error, the time variation of the actual geometric center distance of the driven and driving gears will also cause the change in the tooth backlash. As shown in Fig. 9, the change in the backlash caused by center distance change can be expressed as tan\u0394 = \u0394b a \u03c8 , (45) where \u2206a is the change value of geometric center distance" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002534_012146-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002534_012146-Figure5-1.png", "caption": "Figure 5. Distribution of the magnetic field of cross-section of the Stack 1", "texts": [ " The results of the magnetic field distribution modeling are shown in figures 5-13. To visualize the flow paths of the main magnetic flux: consisting of the sum of the fluxes of permanent magnets and the axial magnetic flux of the superconducting field coil, the simulation results are shown in figures 5 through 13. The figures 5-13 shows that the value of magnetic induction in ferromagnetic parts (tooth, yoke) does not exceed the permissible values. The figures 5 through 8 show cross sections of an electric machine in the plane of the first packet (Fig. 5), in the plane of the axial interpacket yoke (Fig. 6), the second packet (Fig. 7), and also between the second packet and the axial excitation coil (Fig. 8). 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10.1088/1742-6596/1559/1/012146 An analysis of the magnitude of the magnetic induction shows that the iron is not in saturation mode - the magnitude of the induction does not exceed 2.1 T. The direction of magnetization of the poles of 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure10-1.png", "caption": "Figure 10. FE model of a disk-pad assembly", "texts": [], "surrounding_texts": [ "The numerical simulations using the ANSYS finite element software package were performed in this study for a simplified version of a disc brake system which consists of the two main components contributing to squeal the disc and the pads. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. a) Boundary conditions applied to the disk In this FE model, the disk is rigidly constrained at the bolt holes. The bolt holes are tied as rigid body to a reference point, where the rotation of the disk is allowed in the y-axis, its angular velocity is imposed and constant \u03c9= 157,89 rad/s. The support is applied to the hole of the disk and is of cylindrical type of which the degrees of freedom are shown in Matrix 1. b) Boundary conditions and loading applied to the pads Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. The boundary conditions applied to the pads are defined according to the movements authorized by the caliper. Indeed, one of the roles of the caliper is to retain the pads which have the tendency natural to follow the movement of the disk when the two structures are in contact. The caliper maintains also the plates in direction Z. Thus, the conditions imposed on the pads are: \u2022 The pad is embedded on its edges in the orthogonal plan on the contac surface, thus authorizing a rigid movement of the body in the normal direction with the contact such as one can find it in an automobile assembly of brake (Coudeyras;2009) \u2022 A fixed support in the finger pad. \u2022 A pressure P of 1 MPa applied to the piston pad. \u2022 The contact between the pads and the disk has a coefficient of friction is equal to 0.2 . Boundary conditions in embedded configurations are imposed on the models (disc-pad) as shown in Fig. 12 (a) for applying pressure on one side of the pad and Fig.12 (b) for applying pressure on both sides of the pad. Matrix 1. 1 Radial Free Axial Fixed Tangential Fixed Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited." ] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure3.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure3.9-1.png", "caption": "Fig. 3.9 The gripper used in the experiments. a Image of the parallel jaw gripper. b Complex CAD model used for point cloud based collision analysis. c Simplified model, sufficient and used for the KGF concept", "texts": [ " Whenever both surfaces are located below the scan, no clear evaluation can be done for that pixel because the gripper could be in occluded free space or in occluded collisions. The resulting unclassifiable volume is called threat volume (see Fig. 3.8). Collision and threat volume result in an overall collision volume Vc = \u2211 Part \u2211 i sx sydiwi . (3.12) Here, sx and sy are the pixel dimensions, di is the height difference of gripper surface and scan surface in the colliding pixel i and wi is a weight which can be chosen to distinguish between collision and threat volume. Furthermore,wi can give different weights to different parts of the gripper (Fig. 3.9). At this point, it was shown how single grasp poses can be efficiently evaluated. To further enhance the performance, the approach is extended to work on predefined grasp regions. To achieve this, the Key Grasp Frame (KGF) is introduced by the author. A KGF consists of a base pose for the end-effector defined in the CAD 26 3 3D Point Cloud Based Pose Estimation model\u2019s coordinate system, a set of degrees of freedom (DOF) and an associated range for each DOF in which this base pose can be varied", " The gripper images are shifted pixel-wise along the axis of the free DOF over the rendered scan image. Then, at each step, the rendered images are analyzed pixel wise for collisions as explained above. The collision volume is stored for each step resulting in a 1D collision function dependent of the variable parameter. Using this function, all parameter values below a predefined collision threshold tc are located and used to build a distance map for all possible collisions (Fig. 3.11e). If the gripper can bemodeled as cuboids (Fig. 3.9), the described procedure can efficiently be implemented using integral images introduced in [92]. If not, efficient FFT based correlation can be applied to reduce computation times. In case of a rotationalDOFapreprocessing step is needed.To convert the rotational problem into a translational one, the rendered images are transformed into polar coordinates, which transforms the rotation into a shift along the orientation axis (Fig. 3.11d). The subsequent steps are then the same as before. After the parameters for all defined KGFs are computed, the one with the highest distance to all collisions is determined using the calculated distance map (Fig", " To revise, a gripper pose WPG is a pose of the gripper relative to the world coordinate system, at which an object can be grasped. Using depth images as model for the workspace, this means that a gripper pose is a feature in the depth image that shows a local decrease in depth, therefore a local increase in object height. In other words, a pattern must be found in the image which matches the gripper footprint in its appearance. The gripper used in the experiments is a standard parallel jaw gripper that was already used in former experiments (see Fig. 3.9). The footprint of the gripper fingers, which are the important part of it for grasping, can be approximated as two simple squares. As the gripper may be rotated around its approach vector and the algorithm shall be as fast as possible, a filter kernel is used that is rotationally symmetric and thus contains all possible gripper orientations (see Fig. 4.1). This simplification ensures a gripper pose estimation via a single convolution operation. Such a filter kernel can also be used for other gripper types" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001698_icelmach.2014.6960379-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001698_icelmach.2014.6960379-Figure2-1.png", "caption": "Fig. 2. Object of study \u2013 three-phase permanent magnet generator.", "texts": [ " 413/L-4/2012 named \u201cVibroacoustic diagnostic method of traction permanent magnets motors and generators based on the own signals\u201d realized in Institute of Electrical Drives and Machines KOMEL. In order to understand the phenomenom of voltage generated under the influence of machine vibration and to determine the impact of vibrations on the shape of generated voltage, series of computer simulations using FEM software have been carried out. To confirm the validity of calculations, laboratory measurements have been conducted. As an object of study low-voltage three-phase permanent magnet generator has been choosen \u2013 presented in Fig. 2. The most important parameters and design data of PM generator: \u2212 Nominal power: PN = 450 W, \u2212 Rated speed: nN = 750 rpm, \u2212 Nominal voltage: UN = 40 V, \u2212 Number of poles: 2p = 8, \u2212 Shaft height: H = 90 mm, \u2212 Winding: three-phase star-connected, distributed, \u2212 Position of magnets: surface mounted. Analysis of PMSM Vibrations Based on Back-EMF Measurements M. Bara\u0144ski, T. Jarek O 978-1-4799-4389-0/14/$31.00 \u00a92014 IEEE 1492 III. COMPUTER CALCULATIONS AND LABORATORY MEASUREMENTS Computational model of permanent magnet generator has been created in ANSYS Maxwell software" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003453_s12206-021-0334-5-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003453_s12206-021-0334-5-Figure11-1.png", "caption": "Fig. 11. Experimental setup of two mechanisms: compliant five bar and compliant closed chain.", "texts": [ " Relating the lateral and axial tip coordinates of the proposed fully compliant, closed chain mechanism as depicted in Fig. 2(a) to the rigid link angles yields two equations with three unknowns. Since the number of unknowns is more than the number of equations from the geometric constraints, problem can\u2019t be solved. On the other hand, for the initial compliant five bar design shown in Fig. 2(b), rigid links on each side of the mechanism can be calculated using the analytical kinematic analysis along with the initial angular velocities. We created two experimental setups as seen in Fig. 11 and applied the same inputs to the servos and observed that the tip follows the same trajectory if the horizontal distance between actuators were adjusted since the overall geometry of both mechanisms is same. Therefore, we adopted the simplified model of the compliant five bar mechanism [25] to control the tip of the closed chain, fully compliant mechanism for the ease of fast response in real time. Controlling the motion of the system in real-time is accomplished by mimicking drawings created in a drawing app on a tablet as they are made" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002384_s12206-020-0325-y-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002384_s12206-020-0325-y-Figure2-1.png", "caption": "Fig. 2. Bulgy and hollow shapes at the interface between steel and aramid/phenol composite parts.", "texts": [ " Several hybrid gears were successfully produced by the wet laid manufacturing with a specially designed hot-pressing mold and one hybrid gear is shown in Fig. 1. The gear dimension is as follow: r1 (steel shaft inner radius) = 18.5 mm, r2 (steel shaft outer radius) = 21.5 mm, r3 (composite outer radius) = 34.4 mm, r4 (dedendum radius) = 36.5 mm, r5 (addendum radius) = 41.05 mm, and the width of a gear, t = 13.2 mm, as shown in Fig. 1. To secure the interface adhesion between steel and aramid/phenol composite, the shapes of interface are complicated and in bulgy and hollow contour, as shown in Fig. 2. A torsion test was performed for the hybrid gear with an initial interface feature in Fig. 5 in order to determine the failure torque and to investigate the interface failure mode between the composite and steel parts. Torque increases gradually through the central shaft of a hybrid gear until torque failure occurs while the teeth region is fixed. When a torque reached 857 Nm, bearing (compressive) failure was occurred at the contacting interface surface of an aramid/phenol composite part with the inner steel counterpart, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003774_s11431-021-1836-7-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003774_s11431-021-1836-7-Figure1-1.png", "caption": "Figure 1 (Color online) The smallest circumscribed sphere of a quadrotor.", "texts": [ " In Section 2.1, the quadrotor and obstacles are approximated as spheres and directed planes to facilitate collision detection. In Section 2.2, the path planning problem is formed as a nonconvex optimization problem, which is solved in Section 2.3 by SCP. For the efficiency of collision detection, the quadrotor and obstacles are approximated as geometric elements with simple shapes. For accuracy, the quadrotor is approximated as its smallest circumscribed sphere Sq with a radius of r as shown in Figure 1. According to the shapes, the obstacles are approximated as unions of spheres S i n, = 1, 2, \u2026 ,o i s, and directed planes P j n, = 1, 2, \u2026 ,j p. For example, as shown in Figure 2, a cube is approximated as some spheres. An effective method for approximation with spheres is given in ref. [26]. Note that every sphere So i, can be described by its sphere center c i and radius ro i, , and every directed plane Pj can be described by its unit positive normal vector np j, and a parameterDj so that every point Rx 3 in the plane satisfies the equation Dn x + = 0p j j, T " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001483_0954406215612814-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001483_0954406215612814-Figure5-1.png", "caption": "Figure 5. Application of Aronhold\u2019s first theorem: involute tooth profile and return circle.", "texts": [ " In fact, Aronhold\u2019s first theorem states that the return circle is the locus of the centers of curvature of all envelopes whose generating curves are lines; moreover, when a line in the moving plane always passes through a fixed point by sliding and rotating about it, that point is on the return circle. Conversely, Aronhold\u2019s second theorem states that the inflection circle is the locus of the centers of curvature for all generating curves whose envelopes have an inflection point. Thus, referring to Figure 5, is the center of curvature for all envelopes whose generating curves are straight lines, as line , which is attached to the auxiliary straight line centrode \". Consequently, the diameter of the return circle R is always equal to the radius of curvature rl of the gear pitch curve l when the Camus theorem is applied in the form of the at UNIV CALIFORNIA SAN DIEGO on December 14, 2015pic.sagepub.comDownloaded from rack-cutter method. This diameter is constant for circular gears, while it varies according to the evolute of the pitch curve in the case of non-circular gears. However, R is always tangent in P0 and internal to l. When passes through the return pole R, as in the case of 0, its envelope shows a cusp onR, as shown in Figure 5 along with the envelope curve 0 of 0. One of the most important contributions to the kinematics of mechanisms as based on the use of both the return and the inflection circles, is due to Lorenzo Allievi (1856, Milan \u2013 1941, Rome), an Italian engineer best known for his studies on the water-hammer problem, but also interested in the kinematics of planar mechanisms. Allievi made significant contributions to curvature analysis and, in particular, to the analysis of the inflection and the return circles, which were called De La Hire circles by Burmester,12 along with the direct and inverse Ball points, which were defined in Italian as \u2018\u2018ondulazione\u2019\u2019 and \u2018\u2018cuspidazione\u2019\u2019, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003001_j.engfailanal.2020.105195-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003001_j.engfailanal.2020.105195-Figure12-1.png", "caption": "Fig. 12. Bicycle frame loading diagram for the fatigue test with the force applied to: a) pedals F1, b) fork axis (horizontal forces F2, F3), c) saddle (vertical force F4), unit in mm.", "texts": [ " The conditions of the numerical analysis conform to ISO 4210\u20136 [5], developed by a special technical committee TC149 on the standardization in the field of cycles, their components and accessories with particular reference to terminology, testing methods and requirements for performance and safety. The standard includes safety requirements and test methods for bicycle frames for selected loading cases depending on the bicycle type (city and trekking, young adult, mountain, racing). The scope of three analyses for the carried out fatigue tests was determined. Fig. 12 shows the boundary conditions for the analysed cases. The removal of the degrees of freedom was determined in relation to the system axis for offset (Ux,y,z) and rotation (Rx,y,z). The bold lines show a simplified frame geometry corresponding to the actual test object. The loading conditions specified in the standard corresponding to the bicycle user weight of 80 kg. The first case applies to a constant amplitude loading due to force F1 applied to the pedal axis, directed downwards at an angle q (Fig. 12a). For a mountain bicycle, the standard specifies the force F1 = 1200 N. A single loading cycle includes applying and releasing the force from each bicycle pedal. The recommended fatigue life is 100,000 cycles. Due to a symmetrical frame loading, the analyses will be carried out for the loading of a single pedal, corresponding to half of the recommended number of cycles. Angle q in accordance with the standard is 7.5\u25e6 \u00b1 0.5\u25e6. The conditions used in the standard always assume a vertical frame position during loading", " The conditions are consistent with the standard operating conditions of the bicycle frame. In the case of the analyses, the bicycle frame is used in gravity mountain biking where the bicycle frame can deviate from its vertical position. To extend the range of analyses and adapt the load to the actual conditions, the calculations for variable angle (q = 7.5; 15; 22.5; 30\u25e6) were proposed. The other parameters remain constant. The second case apply to a constant amplitude loading for the maximum force (F2) and minimum force (F3). The loading applied horizontally to the fork axis (Fig. 12b). For mountain bicycle, force F2 directed forward is 1200 N and force F3 directed backward is 600 N. The specified forces determine the maximum values for a single stress amplitude cycle. The recommended fatigue life is 50,000 cycles. The third case apply to a constant amplitude loading due to force F4 applied vertically downwards in the seat installation point. The standard specifies force F4 for the mountain bicycle of 1200 N. Fig. 12c shows the dimensions for the seat bracket mounting depth (75 mm) and maximum seat height (250 mm). The values depend on the bicycle type. To correlate the analysis results with the standard, the parameters will not be modified. The recommended fatigue life is 50,000 cycles. T. Tomaszewski Engineering Failure Analysis 122 (2021) 105195 The numerical model was based on the surface model divided into almost 190 thousand finite elements. Quadrilateral elements (SHELL181) with six degrees of freedom in each node (translation and rotation in the direction x, y, z) were used", " Tomaszewski Engineering Failure Analysis 122 (2021) 105195 Table 4 shows the maximum stress for all loading cases, relative stress (Smax/Re) and recommended fatigue life. In other loading cases (Fig. 13b, c), the maximum stress was significantly below the material\u2019s yield strength. Fig. 14 shows the local stress distribution taking into account the geometry of the welded joint at the fatigue crack location. The boundary conditions are consistent with the first case for the force F1 applied to the pedal axis, directed downwards at an angle 7.5\u25e6 (Fig. 12a). The stress distribution along the tube is schematically shown below the distribution map. The maximum stress is located on the inside of the top tube. The stress concentration resulting from the weld geometry does not significantly affect the obtained results because the maximum values are on the opposite side. The actual position of the crack is shown in the Fig. 14. The crack is situated away from the edge of the weld and is close to the extreme of the function. The stress distribution is consistent with the fatigue crack initiation point (Section 2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001066_1.4878629-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001066_1.4878629-Figure3-1.png", "caption": "FIG. 3. Dimensions of the spherical body.", "texts": [ " In this paper, we propose a method for measuring the impact force of a spherical body dropping onto a water surface by modifying the LMM, and its validity is experimentally demonstrated. a)Author to whom correspondence should be addressed. Electronic mail: ryosuke.araki.tech@gmail.com Figure 1 shows a schematic diagram of the experimental setup for evaluating the impact force acting on a spherical body that drops onto a water surface. Figure 2 shows the photographs of the spherical body used in this experiment, and Fig. 3 shows the details of the body itself. The main part of the body is fabricated by machining a spherical SUS440 stainless steel body approximately 30.2 mm in diameter, the surface of which is tempered. A cube corner prism, 12.7 mm in diameter, is inserted with an adhesive agent so that its optical center coincides with the center of gravity of the whole body. The total mass of the body, M, is approximately 93.88 g. The origins of the time and position axes are set to be the time and the position at which the impact force is detected in the water test" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000019_iros.2018.8593927-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000019_iros.2018.8593927-Figure5-1.png", "caption": "Fig. 5: Snapshots (left) and trajectory from two perspectives (right) of dual-arm motion with closed kinematics in simulation.", "texts": [ " The planning times of TrajOpt and GRAMPC scale linearly, whereas the required time for CHOMP heavily increases, because the update (23) involves costly matrix inversions. As shown in [16], a multigrid approach can speed-up planning in these cases. Furthermore, the total number of solved problems decreases for TrajOpt and CHOMP, if Nhor is increased, which indicates that the parameters of the planners can not be chosen independent from the trajectory discretization. In contrast, GRAMPC can solve the same amount of problems without adjusting any parameters. Figure 5 shows an illustrative example of a point-to-point motion with closed kinematic chain, whereby the robot has to avoid collisions with the cylinder-shaped obstacle. This demonstrates that the proposed planner is able to compute trajectories for dual-arm robots with closed kinematics. This paper presented a new optimization-based approach for motion planning of dual-arm robots with closed kinematics. The task is formulated as dynamic optimization problem with constraints for collision avoidance and the closed chain" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure4.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0001375_s10853-014-8124-4-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001375_s10853-014-8124-4-Figure5-1.png", "caption": "Fig. 5 Possible growth inhibition mechanism of strain E. coli with CNTs", "texts": [ " The ROS is an another factor, which is produced in the bacteria solution with F-MWCNTs and are responsible to form the free radicals in the solution, and these free radicals penetrate the outer wall of the cells and enter to inner wall of the membrane. These free radicals when react with the cytoplasm of the bacteria, enzymatic changes occurred and it leads to the disorganization of the cell leakage of the cell contents. The exact mechanism of membrane damage is not clear and it is still under debate. Further conclusive studies are needed to conclude the relation of antibacterial activity with F-MWCNTs. Moreover, on the basis of their obtained observations, we have proposed a possible mechanism (Fig. 5). Most likely in all the pathogenic strains tested with F-MWCNTs have attached initially to the outer membrane (Fig. 5a) of the cell resulting in formation of pits (Fig. 5b) in the cell wall followed by breakdown of the outer membrane of the cell (Fig. 5c). This has led to morphological change in the cell resulting in cell lyses (Fig. 5d). Although, the biochemical studies, which are in due course are needed to conclude the exact relation between growth inhibition activity of bacteria with the prepared F-MWCNTs, but as the observation, the prepared structure is a potential candidate to inhibit the growth of bacteria. Analytical determinations of functionalized CNTs for the inhibition of bacterial growth The determination of concentration of functionalized multiwall carbon nanotubes (F-MWCNTs) for the inhibition of bacteria growth was studied by analytical methods" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002322_s11665-020-04735-8-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002322_s11665-020-04735-8-Figure3-1.png", "caption": "Fig. 3 Heat transfer model of a continuously moving point heat source", "texts": [ " T A; t\u00f0 \u00de \u00bc r t 0 dTA \u00f0Eq 1\u00de T \u00bc 2Q cq 4pat\u00f0 \u00de3=2 exp R2 4at \u00f0Eq 2\u00de Here, the nomenclature used is as follows: T\u2014temperature as a function of time, t and distance, R; Q\u2014the instantaneous heat energy (J); R\u2014distance from the heat source, R \u00bc x2 \u00fe y2 \u00fe z2\u00f0 \u00de1=2 (mm); t\u2014 time (s); a\u2014thermal diffusivity of material (m2/s); q\u2014density of material (kg/ m3); c\u2014specific heat capacity of material [J/(kg C)]; k\u2014thermal conductivity of material (W/m C). T \u00bc r t 0 2qdt0 cq 4pa t t0\u00f0 \u00de\u00bd 3=2 exp R02 4a t t0\u00f0 \u00de \u00f0Eq 3\u00de where R02 \u00bc x0 vt0\u00f0 \u00de2\u00fey20 \u00fe z20; q\u2014the effective thermal power of a heat source (W); v\u2014the laser scanning speed (mm/s). Assuming that the initial temperature of the work-piece is 0 C, surface heat dissipation is not considered at the same time. That is, when the heat source is at O\u00a2, the instantaneous coordinate distance to point A is R\u00a2, as shown in Fig. 3. In order to solve the problem, the mobile coordinate is adopted, that is, the position of the heat source is taken as the origin: T x;y;z; t\u00f0 \u00de\u00bc 2q cq 4pa\u00f0 \u00de3=2 exp vx 2a r t 0 dt00 t003=2 exp v2t00 4a R2 4at00 \u00f0Eq 4\u00de Journal of Materials Engineering and Performance where R2 \u00bc x2 \u00fe y2 \u00fe z2; x \u00bc x0 vt; y \u00bc y0; z \u00bc z0; t00 \u00bc t t0; when t fi \u00a5, v = constant, q = constant; u2 \u00bc R2 4at00 ;m 2 \u00bc R2v2 16a2 ; Also, r 1 0 e u2 m2 u2 du \u00bc ffiffi p p 2 e 2m, Therefore, T \u00bc q 2pkR exp vx 2a Rv 2a \u00f0Eq 5\u00de The solidus of 40Cr steel is between 1409 and 1429 C, and the liquidus of 40Cr steel is between 1486 and 1504 C" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002286_j.cagd.2020.101826-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002286_j.cagd.2020.101826-Figure8-1.png", "caption": "Fig. 8. Depth interposition: anamorphic deformation of a straight bar S .", "texts": [ " This direct 3D printing represents a substantial improvement over simply pasting 2D printed textures on the model, as done by Tsuruno (2015), who observe that their option does not apply to sculptured surfaces. Next, we show detailed examples of how to materialize impossible objects, via anamorphic deformation of textured NURBS. Most generic views of an object S , requiring hidden-line removal from a given viewpoint O, can be turned into an impossible figure. To achieve an occlusion paradox, simply apply a local anamorphic deformation, on a hidden part of S , so that it becomes visible. In the example of Fig. 8, we defined each face of the horizontal bar in S with three consecutive planar facets and then bent the central ones. These central facets must be refined, as explained in Section 3.3, to introduce additional degrees of freedom for the desired effect, namely altering the depth ordering between the bar and the central vertical column. The tile texture, aligned with the original edges, help mask the deformation and the resulting shading artifact. In this figure (and also in the upcoming Fig. 9 and Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003664_tia.2021.3084549-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003664_tia.2021.3084549-Figure9-1.png", "caption": "Fig. 9. Distribution of flux density at cross-section view (3D-FEA result).", "texts": [ "org/publications_standards/publications/rights/index.html for more information. in a conventional axial gap motor that is not consequent-polePM-type. However, the proposed motor allows to ignore the effects of the even-order components which cause asymmetric voltage amplitude and reduce the voltage utilization of the inverter. Table III shows the winding factor of the proposed motor which uses combination of 20 poles and 24 slots, that of even-orders are 0. Therefore, the harmonic components of even-orders can be canceled [16]. Fig. 9 shows the magnetic flux density distribution of the cross-section by 3D-FEA under no field current and under maximum field strengthening. At maximum field strengthening, the magnetic flux density at the shaft, the case lid, and the SMC is higher than under no field current. The SMC is effectively excited by the field winding. The thickness of the PM is designed to have sufficient demagnetization resistance in the proposed motor, although demagnetization is a concern because the magnetic field is generated in the opposite direction to the PM's magnetization under field strengthening" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002682_j.procir.2020.02.209-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002682_j.procir.2020.02.209-Figure4-1.png", "caption": "Fig. 4. Structural error energy in mJ on a logarithmic scale for Cartesian (a), Cartesian with projection factor (b) and layered tetrahedral (c) meshing with size 0.5 mm.", "texts": [ " Another parameter that has to be set on the L-PBF machine and which is not present in simulations is the air speed of the inert gas. In this case it was set to 2.5 m/s, resulting in an inert gas flow of 4.4 l/min. In order to assure a non-oxidizing build environment, the maximum level of residual oxygen in the build chamber was allowed to be 0.01%. At the beginning of the print the temperature inside the build chamber was 24.5 \u00b0C with a relative humidity of 40%. The simulation results for the structural error energy (SERR) and the displacement in z-direction are shown in figure 4 and 5. These results are only two of the result values and model properties that were used to compare the three mesh types. An overview of the values for each mesh type and mesh size is listed in Table 1. Starting with the number of nodes it can be seen that for each mesh size the layered tetrahedron mesh has up to almost double as much nodes as the Cartesian meshes. For the number of elements, this difference is even bigger. This leads to a higher number of total degrees of freedom in the model and therefore to more equations to be solved", " Mesh statistics show, as Fig. 3. TiAl6V4 test sample produced using the L-PBF process. expected, the best value of the average element quality for the Cartesian mesh with 99.4%. Even the elements with distortion induced by the projection factor have still an average quality above 90%. This is also reflected by the low minimum element quality of the tetrahedral mesh with approximately 20%. As the previous mentioned values and properties only give an overview on the simulation process itself, the results of figure 4 and 5 are used to compare the mesh types for result accuracy. In both colour plots can be seen that the inconsistent colour variations, especially for the Cartesian mesh, are dependent on geometry representation accuracy. The fast and inhomogeneous colour transitions are an indicator for inaccurate results. This can also be seen in figure 4 as it shows the estimated structural error energy (SERR) on a logarithmic scale with up to the order of two higher results for the Cartesian mesh. In fact the arithmetic mean value over all elements in the model is 0.3 mJ for the layered tetrahedral mesh compared to 2.1 mJ for the Cartesian. This difference in result accuracy is not changing with even smaller elements as all models reached result convergence. The layered tetrahedron mesh converged even already with a mesh size of 0.75 mm. As criteria for a converged model the gradient of the estimated error has to be of an order close to the squared characteristic element length following the example of [13] for displacement errors", " The red color indicates a positive deviation from the simulation geometry, which means the diameter of the sample is larger than the simulation. This is consistent with a total shrinkage of the sample in z-direction. Deviations on the cylindrical parts of the test sample should be neglected as they are not constant in x- and y-direction and hence afflicted with inaccuracy in alignment of both z-axis. The result accuracy shows a strong dependency on the geometry representation, especially seen in figure 4, at each point of the mesh. In other words, Cartesian meshing is still reasonable for models with less curvature. Nevertheless the tetrahedral mesh can be coarser and showed, other than the Cartesian meshes, also with 0.75 mm element size, accurate results. For almost the same accuracy the 0.75 mm tetrahedral mesh needed only 3,481s. The achieved time savings due to reduced computational cost are in the order of ten. The layered tetrahedral mesh shows a more geometrical accurate and therefore more realistic representation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001152_s12206-014-0814-y-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001152_s12206-014-0814-y-Figure1-1.png", "caption": "Fig. 1. The rubbing experiment of the rotor experiment rig of aeroengine.", "texts": [ " Firstly, its shape is consistent with core-engine's casing, and its size is treble reduced. Secondly, the internal structure is simplified, that is the core-engine is simplified to 0-2-0 support structure form and adjustable stiffness support structure is designed to adjust system's dynamic characteristics. Thirdly, multistage compressor is simplified to singlestage disk structure. Finally the experimental rig of aero-engine forms the rotor-support-blade disk -casing system in structure. The experiment rig is shown in Fig. 1. Four rubbing screws are designed on turbine casing to carry out the rubbing experiments at the four positions. Four accelerometers are placed on the turbine casing to collect the casing acceleration signals. The rubbing positions and the installation positions of accelerometers are shown in Figs. 1 and 2. Facing to the turbine casing, the directions of rubbing positions and installation positions of accelerometers are shown in Fig. 2, which also shows the testing channels of the four accelerometers" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000871_12.2185196-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000871_12.2185196-Figure3-1.png", "caption": "Figure 3. Force and moment balance applied on a quadrotor.", "texts": [ " Consider a ZYX rotation matrix to express the attitude of the body frame with respect to the inertial frame: (1) where c and s refer to cosine and sine, respectively. Proc. of SPIE Vol. 9468 94680R-3 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 07/20/2015 Terms of Use: http://spiedl.org/terms Consider the basic equations of motion by balancing forces and moments. Assume that the center of gravity of the body is positioned in the same plane as the quadrotor arms. This assumption simplifies the equations, although it is not without its drawbacks, as discussed later. Figure 3 depicts principle forces and moments acting on the quadrotor, l is the length of a single rotor arm measured from the center of gravity to the rotor hub. All rotor arms are assumed to have the same length. Based on Fig. 3, the basic equations of motion are defined as follows: \u2211 sin sin cos cos sin \u2211 cos sin sin cos sin \u2211 cos cos \u2044 \u2044 \u2044 (2) where Ji are the moments of inertia with respect to the axes and l represents the length of an individual rotor arm from the axis origin. Ki refer to the drag coefficients. For the purposes of this model, assume that the drag is zero, which may be reasonable at low speeds. Define the controller inputs, as direct functions of the required forces. Furthermore, assume a linear relationship between the moments generated by the motors and the thrusts, defining a force-torque scaling factor C, such that F = C\u03c4" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000709_psc49016.2019.9081455-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000709_psc49016.2019.9081455-Figure9-1.png", "caption": "Fig. 9. motors passive tracks [8]", "texts": [ " Maglev linear motors are divided into two main categorize of induction and synchronous motors. In terms of the position of armature, they are divided into long armature, in which armature is placed on the track, or short armature in which armature is placed under the train [8]. This categorization is illustrated in Fig. 8. In this section, different types of Maglev motors are explained. A. Linear Synchronous motors Linear synchronous machines have either a passive or an active track. In motors passive tracks, as shown in Fig. 9, both of the field and armature windings are placed in the train and by changing the reluctance of the track, polarity of the motor\u2019s flux is changed and the train is moved [8]. In a similar topology, field windings are replaced with PMs, as shown in Fig. 10. The PMs are placed between the armature windings and in the magnetic loop [9]. In order to improve the performance of the motor under fault conditions, one must reduce the mutual inductance of phase windings with each other. To that end, the design of Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001937_j.optlastec.2016.08.011-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001937_j.optlastec.2016.08.011-Figure4-1.png", "caption": "Fig. 4. Schematic of offset of droplet centroid.", "texts": [ " The welding parameters are shown in Table 2. The droplet transfer phenomena in hybrid welding processes were observed and investigated using a high speed video camera (CA-D6-0256W) with the maximum recording speed of 1000 frames per second. In order to characterize the stability of droplet transfer, offset of droplet centroid Lm and critical additional axial acceleration ac are adopted in this paper. Offset of droplet centroid Lm, which means the vertical distance between droplet centroid and the extension line of the wire axis (Fig. 4), is used to characterize the axial transfer properties of droplets. Generally, less offset of droplet centroid indicates better axial transfer properties of droplets and more stable droplet transfer. Critical additional axial acceleration is used to characterize transfer frequency. The additional axial acceleration is produced by an assumed additional mechanical force that acts on droplets and points to the molten pool along the wire axis. This additional force generates mechanical potential energy and the total energy of droplets increases" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001821_s0263574716000345-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001821_s0263574716000345-Figure3-1.png", "caption": "Fig. 3. Two similar 3-UPU manipulators with different sizes.", "texts": [ " (6) yields \u03c7 = \u221a det ( MTM ) I\u220f i=1 \u2225\u2225$i \u2225\u2225 \u00b7 J\u220f j=1 \u221a det ( MT j Mj ) = \u221a det ( MT 0 M0 ) I\u220f i=1 \u2225\u2225$i \u2225\u2225 \u00b7 J\u220f j=1 \u221a det ( MT j0Mj0 ) . (10) The above equation indicates that the proposed metric of closeness to singularities is invariant with respect to the selection of base screws for the overconstrained or lower mobility parallel manipulator with indefinite base screws. The similar manipulators are defined as the mechanisms with the same architecture and the same length ratio of corresponding links, but with different sizes, as shown in Fig. 3. When two similar manipulators are at the same posture, it is reasonable to expect that the closeness to singularities of these two manipulators should be the same; however, almost neither existing metrics can obtain the identical result, except the method of characteristic angles proposed by Bu.29 The weight factor l proposed in Section 2 can be regarded as a property of the manipulator, which may be defined as the mean or maximum value of distances from the origin of the coordinate frame to axes of all the revolute joints. Under this definition, this weight factor varies as the manipulator moves to different configurations. Alternatively, the weight factor l may be defined as a constant, for example, the length of some link in the manipulator, or the height of the manipulator at its initial configuration. There are two similar manipulators A and B in Fig. 3, and the length ratio of the corresponding links in two manipulators is \u03bb. Suppose $Ai and $Bi (i = 1, 2, . . . , N) denote the corresponding screws of two similar manipulators respectively, and for each manipulator there are p (0 \u2264 p \u2264 N) http://dx.doi.org/10.1017/S0263574716000345 Downloaded from http:/www.cambridge.org/core. New York University Libraries, on 10 Dec 2016 at 01:23:18, subject to the Cambridge Core terms of use, available at http:/www.cambridge.org/core/terms. screws whose real parts are 0", " For the manifold spanned by J submanifolds whose base screws are indefinite, there are J -1 characteristic angles between these submanifolds, but there are no characteristic angles within each submanifold since its base screws are indefinite. Hence, for the entire manifold consisting of I definite screws and J submanifolds, there are I + J -1 characteristic angles between these screws and submanifolds. Since the possible value of the geometric average normalized volume ranges from 0 to 1, the average characteristic angle can be defined as follows. \u03d5 = arcsin (\u03be ) . (15) Two similar 3-UPU parallel manipulators are shown in Fig. 3. The moving platform of each manipulator is connected to the base with three limbs. Each limb consists of two universal joints and a prismatic joint equipped with an actuator. The structural parameters of the larger manipulator are as follows. The three universal joints on the base form an equilateral triangle with the side length 0.5 m, and the origin O of the base frame is at the center of this triangle. Similarly, the three universal joints on the moving platform form an equilateral triangle with the side length 0", " New York University Libraries, on 10 Dec 2016 at 01:23:18, subject to the Cambridge Core terms of use, available at http:/www.cambridge.org/core/terms. Note that the 3-UPU parallel manipulator is a lower mobility mechanism, which indicates the limb constraint wrenches are indefinite when its actuators are locked. Nevertheless, the proposed metric of the geometric average normalized volume or the average characteristic angle can solve this problem. Furthermore, for the smaller 3-UPU manipulator in Fig. 3, the proposed metric can obtain the same result when the smaller manipulator moves along the similar trajectory. http://dx.doi.org/10.1017/S0263574716000345 Downloaded from http:/www.cambridge.org/core. New York University Libraries, on 10 Dec 2016 at 01:23:18, subject to the Cambridge Core terms of use, available at http:/www.cambridge.org/core/terms. Manipulators should be prevented from reaching singular configurations, and the singularity margin that indicates the \u201cdistance\u201d between the current configuration and a singular configuration should be measured" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure13.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure13.2-1.png", "caption": "Figure 13.2 Socket screw heads", "texts": [ "2 mm) steel) or other thin materials and are then screwed home in a single operation. Bolts are used in conjunction with a nut for heavier applications than screws. Unlike screws, bolts are variety of head drives as outlined on page 39 with Phillips, Pozidriv and Torx the most common. Manufactured in high-grade alloy steel with rolled threads, this type of screw is used for higher strength applications than machine screws. Three head shapes are available, all of which contain a hexagon socket for tightening and loosening using a hexagon key, Fig.\u00a013.2. Headless screws of this type \u2013 known as socket set screws \u2013 are available with different shapes of point. These are used like grub screws, where space is limited, but for higher strength applications. Different points are used either to bite into the metal surface to prevent loosening or, in the case of a dog point, to tighten without damage to the work, Fig.\u00a013.3. Self-tapping screws are used for fast-assembly work. They also offer good resistance to loosening through vibration. These screws are specially hardened and produce their own threads as they are screwed into a prepared pilot hole, thus eliminating the need for a separate tapping operation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001918_978-3-319-13117-7-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001918_978-3-319-13117-7-Figure11-1.png", "caption": "Fig. 11. Picture of the flexor and rotational mechanism during simulation.", "texts": [ " SIMULATION OF THE EXOSKELETON MECHANISM SimMechanics MatLab software module to simulate the kinematic behavior of the mechanisms from control laws for the trajectory tracking designed, it was used. It provides a simulation environment for 3D mechanical multibody systems, from the models designed in SOLIDWORKS. Each of the models was translated by the program as a system block diagrams, which showed each component and junction that contains the mechanism involved in the movement. The flexor mechanism is shown in Fig. 9 and the rotation in Fig. 10. Following that, the behavior of the links was simulated, as shown in the images of Fig. 11. Fig. 10. Block diagram of the rotation mechanism. Using Simulink module, it was created a trajectory generator block, which allowed applying control laws in DYNAMIXEL AX-12W actuators. These actuators have been selected because their features allow monitoring the position Exoskeleton for Rehabilitation of Thumb 255 IFMBE Proceedings Vol. 49 of the trajectories of the links of the exoskeleton, to assess how their behavior will be during the course of each therapy. Paths are programmed by the physiotherapist using the user interface according to the record of the patient evolution", " RESULTS It has been defined the ideal geometry for each of the prosthetic components especially for the impulsion system constituted by three carbon fiber springs as well as for a small system to replicate the movements of inversion and eversion located among the superior and the inferior plaques of carbon fiber. According to the above, the three arcs that form the arciform structures of a human foot such as height of the subastragalar joint are capable of supporting a 80 kg pacient. At this point, because of the materials in its composition, the prosthetic device is considered as an active prosthesis of dynamic response (See Figure 9, Figure 10, Figure 11, and Figure 12) which is capable of supporting the FRS of 660 N in the 17% of the gait cycle corresponding to the medium phase of posture, as well as 715 N in the frame 51 of 43 % of the gait cycle corresponding to the terminal phase of posture. For future work, it is proposed the establishment of a manufacturing method for each of the pieces which integrate the impulsion system as well as its analysis through the finite element method by the Pre Post ACP module of Ansys\u00ae. Figure 9 Foot average measures for Mexican users Figure 10 Type C Plaque to absorb the FRS. Figure 11 Superior plaque in carbon fiber for support of pyramidal module and system of inversion and eversi\u00f3n Figure 12 System of movement for inversion and eversion IV. CONCLUSIONS The design of prosthetic systems is without a doubt an engaging field, even more if the design involves a fully identified need with the purpose of replicating the movements of a lost extremity to accomplish a series of requirements that make this design a feasible one to manufacture strengthening the development of projects related to this kind of advices in the country" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001157_cjme.2014.0813.133-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001157_cjme.2014.0813.133-Figure1-1.png", "caption": "Fig. 1. Heat source, geometry, and boundary condition", "texts": [ " With the method and results of this study, the scuffing resistant design can be achieved for tapered roller bearings. The tapered rollers roll and slide on the inner raceway and outer raceway inside the bearing. The heat source model proposed by GECIM, et al[8], is used in present numerical modeling. Furthermore, the tapered roller bearing is divided into ten sections in the generation line of the tapered roller direction. Each slice can be approximated with a cylinder. The temperature distributions of tapered roller and raceways in each section are calculated respectively. Fig. 1 shows the geometry and boundary of the slice model. The heat conductions in axial and circumferential directions are neglected, and the circumferential heat convection is taken into account. The heat flux distribution is uniform over the heating zone, 0 . Because the Hertzian contact width is very small, compared with the circumferential length of the tapered rollers, this assumption is reasonable. With the above assumptions, the equation of heat conduction is 2 2 1 , T T T r rr \u00b6 \u00b6 \u00b6 + = \u00b6 \u00b6\u00b6 (1) where T is the difference between the actual temperature and the ambient ( \u2103 ), r and are the cylindrical coordinates, \u03c9 is angular speed of tapered roller (rad/s), and is thermal diffusivity (m2/s)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003123_msec2015-9240-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003123_msec2015-9240-Figure1-1.png", "caption": "FIGURE 1. ILLUSTRATION OF THE ELLIPSOIDAL HEAT DISTRIBUTION AND THE LOCAL HEAT SOURCE COORDINATE SYSTEM WITH THE ORIGIN AT THE START OF THE HEATING PATH. FOR vs > 0, THE HEAT SOURCE MOVES IN THE LOCAL \u2212z DIRECTION.", "texts": [], "surrounding_texts": [ "If the length of the line input segment is much longer than the heat source radius (i.e. vs\u2206t \u226b c), then the line input is nearly discontinuous and the resulting temperature profiles include nonphysical stairsteps, as shown later in Figure 7(a). These stairsteps consist of regions of approximately zero thermal gradient within each line input segment, and regions of large thermal gradients at the edges of each segment. To smooth out the temperature profile, an elongated ellipse model is proposed where the peak value of the heat input occurs in the middle of the segment and c is on the order of the segment length. To result in a smooth distribution over multiple segments, the power density for the beginning and end of each segment should be half of the peak value. This condition is illustrated in Figure 2 and described by: Qmax exp ( \u2212 3 ( 1 2 vs\u2206t )2 c\u0303 2 ) = 1 2 Qmax (16) from which it follows that the elongated ellipse length c\u0303 is: c\u0303 = vs\u2206t 2 \u221a 3 log2 (17) Then the elongated ellipse power density distribution Q\u0303 is: Q\u0303 = 6 \u221a 3P\u03b7 abc\u0303\u03c0 \u221a \u03c0 exp ( \u22123x2 a2 \u2212 3y2 b2 \u2212 3 ( z+ vs ( t0 + 1 2 \u2206t ))2 c\u0303 2 ) (18) Note that the only differences between Eq. (12) and Eq. (18) are the use of the elongated length c\u0303 in place of c, which stretches the distribution in the local z direction, and the additional 1 2 \u2206t term, which shifts the peak heat input from the end to the middle of the segment. An example of the three power distributions Q, Q, and Q\u0303 compared side by side in the local x-z plane is shown in Figure 3." ] }, { "image_filename": "designv11_22_0002092_0954406216682768-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002092_0954406216682768-Figure1-1.png", "caption": "Figure 1. Ball screw drive.", "texts": [ " The stiffness of the double-nut is tested on the constructed ball screw drive based on the continuous measurement scheme. The contrastive analysis of the experimental results and simulated results shows that the experimental results agree much better with the axial stiffness calculated by the new improved model. Models for the calculation of the axial stiffness of the nut The most common feed drive consists of a ball screw supported by thrust bearings at two ends, and is driven by a servomotor through a coupling, and nuts with recirculating balls.18 A schematic of the ball screw drive is shown in Figure 1. The supporting bearings, the screw, and the ball nut contribute to the axial stiffness of the ball screw drive, and the stiffness is represented with three springs connected in series, as shown in Figure 2.4,13 kbr represents the bearings stiffness, kb represents the ball screw stiffness, and kn represents the equivalent axial stiffness of the preloaded nut. The nut is an important component of the ball screw drive. In order to calculate the relationship between axial contact force and axial deformation of a nut, Hertz contact theory is used to calculate the nominal deformation and contact force" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001755_978-3-319-26327-4-Figure11.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001755_978-3-319-26327-4-Figure11.2-1.png", "caption": "Fig. 11.2 Parrot AR.Drone schematics with and without indoor hull (image based on: http://ardrone2.parrot.com/ ardrone-2/specifications/)", "texts": [ " This is a complex model with a hybrid reference frame, where the translational motion equations are expressed with respect to the world frame and the rotational motion equations are given with respect to the quadcopter body frame. Instead of this, we are going to use a simplemodel, presented in Sect. 11.4.2, which will include only the yaw rotation of the quadcopter combined with a constant velocity model, because we are not focusing on quadcopter flight stabilization control, but on higher level control. The chosen quadcopter is the Parrot AR.Drone 2.0, see Fig. 11.2 for a schematic. It is a low-cost but well-equipped drone suitable for fast development of research applications (Krajn\u00edk et al. 2011). The drone has a 1GHz 32 bit ARM Cortex A8 processor with dedicated video processor at 800 MHz. A Linux operating system is running on the on board micro-controller. The AR.Drone is equipped with an IMU composed of a 3 axis gyroscope with 2000\u25e6/s precision, a 3 axis accelerometer with \u00b150 mg precision, and a 3 axis magnetometer with 6\u25e6 precision. Moreover, it is supplied with two cameras, one at the front of the quadcopter having a wide angle lens, 92\u25e6 diagonal and streaming 720p signal with 30 fps" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001914_1.b35750-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001914_1.b35750-Figure12-1.png", "caption": "Fig. 12 Vortex in labyrinth seal cavities.", "texts": [ " The details of this phenomenon can be analyzed by combining themwith the flow analysis in the next section. Labyrinth seals can hold back leakage with the throttling effect of the seal tooth or by dissipation of the vorticity in the seal cavities. The difference in vortex flows between the two kinds of labyrinth seals is shown in Fig. 11. The vortices are strong and visible in the cavities of the straight-through and interlaced labyrinth seals. According to the distribution of the limiting streamline in Fig. 11, one seal stage is divided into three areas (Fig. 12). A comparison of the two kinds of labyrinth seals shows that the quantities and spacing of the axial teeth tips are identical, and the only distinction is the arrangement of seal teeth. In the straight-through labyrinth seal, all sealing teeth are aligned. Clearance between the tip of the seal and the rotor is immediate, and a portion of leakage gas can flow via passage 1\u20132\u20131 (in Fig. 12a), and not into the seal cavities. In Fig. 12a, vortex flow occurs mainly in the no. 3 area. Therefore, plenty of leakage gas, which did not flow into the cavities, is not blended into the vortex. The dissipation effect of the seal cavities is therefore ineffective on the aforementioned portion of leakage gas. Adjacent teeth are staggered in the interlaced labyrinth seals. Leakage gas, which flows out from the seal teeth, cannot enter the 0 A er od yn am ic f or ce F /( N ) 5 15 10 -15 -10 -5 P1 P2 P3 P4 P5 P6 P7 P8 Rotor position 0 20 -20 P1 P2 P3 P4 P5 P6 P7 P8 5 15 10 -15 -10 -5 A er od yn am ic f or ce F /( N ) Rotor position Amplitude 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001170_0278364914551773-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001170_0278364914551773-Figure6-1.png", "caption": "Fig. 6. Rocky7 Frames. Two views of the left side are shown. Wheels 2, 4, and 6 on the right correspond to 1, 3, and 5 on the left. Figure is inspired from (Tarokh and McDermott, 2005).", "texts": [ " The overall design of the rover is illustrated in Figure 5. The main rocker angle r is antisymmetric, having opposite values on each side of the body on uneven terrain (r1 = 2r2). The two smaller rockers on either side are passive (and independent) at their \u2018\u2018bogie\u2019\u2019 joints b1 and b2. This setup provides exactly the three passive degrees of freedom needed to allow the six wheels to remain in contact with the terrain. Coordinate frames are assigned according to the DH convention as in (Tarokh and McDermott, 2005) in Figure 6. Only the left side is shown. Frames A2, A4, A6, B2, S2, S4, and S6 occur symmetrically on the right side. Likewise, the R frame is positioned relative to some world frame and that transform has six degrees of freedom so the DH convention does not apply. Our RKP approach is consistent with the use of DH (or any other convention) because the important part is to not differentiate the position kinematics. Indeed, the parameters in the DH transform table can be employed directly in the recursive relations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure2-1.png", "caption": "Fig. 2. BESO designs for nodes under bending moment and axial forces.", "texts": [ " BESO parameters, including evolutionary ratio, volume fraction and filter radius, are set to 2%, 95%, and 3 mm respectively. Laplacian smoothing algorithm is then applied to the optimised node to smooth rough surface caused by element removal. The geometry of the node, which is used as the initial model in BESO design, as well as its dimensions are shown in Fig. 1. In this model, the non-design domain consists of small rings around the bolt hole. Two symmetrical load cases applied in the design process include axial compressive loads and out-of-plane bending moment. Fig. 2 shows the rendered perspective views of the nodes designed for symmetrical bending moment and axial compressive forces. By comparing the BESO designs for axial forces (axial node) and out-of-plane bending moment (bending node), it can be clearly observed that the optimised planar topologies for both nodes are almost identical for top and bottom planes. In the axial node, loads are applied on the top and bottom planes, whereas the planar structures in the bending node are formed in planes parallel to the planes of the applied loads" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002534_012146-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002534_012146-Figure4-1.png", "caption": "Figure 4. Three-dimensional design of the HTS generator", "texts": [ " Figure 2 shows that the windings of the armature right packages are rotated by 15 geometric degrees. Three-dimensional sketch of the rotor is shown in figure 3. Rotor packages are made of sheets of electrical steel; permanent magnets are in their closed grooves. The insignificant amount of pole scattering fluxes is compensated by a decrease in the air gap. The hollow shaft has a lower mass and sufficient rigidity to increase the critical speed. A variant of the elaborated principal design is shown in figure 4. 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10.1088/1742-6596/1559/1/012146 . 1 \u2013 axial excitation coil; 2,3 \u2013 ferromagnetic poles; 4 \u2013 shaft; 5 \u2013 permanent magnets; 6 \u2013 armature windings; 7 \u2013 stators yoke; 8 \u2013 stators teeth; 9 \u2013 bandage (ferromagnetic, part of sheetselectrical steel); 10 \u2013 inter-package axial yoke; 11 \u2013 axial yoke; 12 \u2013 one bearing shield; 13 \u2013 bearing shield-shaft radial air gap; 14 \u2013 bearing shield-shaft axial air gap; 15 \u2013 framework; 16 \u2013 reserve excitation coils; 17 \u2013 reserve armature windings; 18 \u2013 stator tube" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002854_j.engfailanal.2020.105028-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002854_j.engfailanal.2020.105028-Figure14-1.png", "caption": "Fig. 14. Safety factor of new scheme.", "texts": [ " In order to ensure the same lubrication conditions, the diameter of the oil passage is increased by 1.4 times. The optimized structure is shown in Fig. 13. It can be seen that the cross-sectional area and transition area of the connecting rod Z. Pan and Y. Zhang Engineering Failure Analysis xxx (xxxx) xxx have been modified, the modified cross-sectional area is increased by about 30%. The fatigue safety factor of key points in the new scheme is analyzed by applying the same load and boundary constraints. From Fig. 14, the fatigue safety factors of points 1 and 2 are greatly improved. After increasing the cross-sectional area of the connecting rod, the fatigue safety factor of the connecting rod is improved. However, the weight of the connecting rod is also increased. With the increase of cross-sectional area, the weight of the connecting rod is increased by 0.8 kg and the power to weight ratio of the engine is reduced by 0.01, which has little influence on the power weight ratio of the whole machine. But the reliability has been greatly improved" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001313_s00419-014-0824-3-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001313_s00419-014-0824-3-Figure1-1.png", "caption": "Fig. 1 2D representation of the bearing (left-hand side)\u2014misalignment contribution (right-hand side) [9]", "texts": [ "2 Basic fluid dynamics: Reynolds equation The dynamic boundaries as well as the surrounding geometric conditions of the oil-film bearing system allow the consideration of certain assumptions and simplifications, which applied to the governing Eqs. (5)\u2013(6) lead to the derivation of the basic Reynolds equation for oil-film lubrication. The assumptions are listed below: 1. Oil-film properties (a) Negligibly small inertia forces compared to the associated friction forces. (b) Constant density \u03c1 and viscosity \u03bc over the whole area of interest. (c) Newtonian fluid undergoing laminar flow. (d) Negligibly small first- or higher-order velocity gradients in x- and z-direction compared to the asso- ciated gradient in y-direction (Fig. 1). (e) Negligibly small pressure variation in y-direction compared to the associated pressure in x- and z-direction (Fig. 1). 2. Geometric properties (a) Ideally smooth and rigid surfaces, but still friction afflicted. (b) Negligibly small clearanced curvatures. (c) Small bearing clearance compared to the associated bearing dimensions. Based on the first assumption 1(a) Eq. (5) is simplified to give Eq. (7), which according to 1(d)\u2013(e) is rewritten as follows (8)\u2013(10): (5) 1(a)\u21d2 \u2207 p \u00b7 I = \u03c1g + \u03bc v, (7) \u03d1p \u03d1x = \u03bc \u03d12u \u03d1y2 (8) \u03d1p \u03d1y = 0 (9) \u03d1p \u03d1z = \u03bc \u03d12w \u03d1y2 + \u03c1g (10) The double integration of (8), (10) provides the following solution for the velocity components u and w, i", " (12) The term h := h(x, z) in (12) refers to the oil-film thickness formulated in (13), and U, V,W are the velocity components of the associated shaft, which due to the presence of friction coincide with the velocity components of the fluid on the bearing\u2019s surface (14). For a detailed description as well as the derivation of (13) and (14) see Appendix A. h(\u03d5, z) = R \u2212 r + e(z)cos(\u03d5 \u2212 \u03c6(z)) and h(\u03d5, z) = l \u2212 z l h(\u03d5)|z=0 + z l h(\u03d5)|z=l (13) [U V W ]T \u2261 [u(x, y, h(x, z)) v(x, y, h(x, z)) w(x, y, h(x, z))]T (14) Equation (12) can be extended by considering the misallignment effect of the shaft/bearing system as depicted in (Fig. 1). Taking into consideration the 2(c) assumption, the following holds: \u03d5 << 1 \u21d2 cos\u03d5 \u223c= 1 \u21d2 W = Wscos\u03d5 \u223c= Ws (15) and thus, the velocity component V is expressed as the sum of the time rate of change for the oil-film thickness and the velocity components U and W , i.e., V = \u03d1h \u03d1 t + W tan\u03c81 + U tan\u03c82 (15)= \u03d1h \u03d1 t + Ws sin\u03c81 + U sin\u03c82 = \u03d1h \u03d1 t + W \u03d1h \u03d1z + U \u03d1h \u03d1x . (16) The substitution of (16) in (12) along with the definition of the oil-film thickness given in (13) delivers the extension of the previously formulated Reynolds equation, which also accounts for the misalignment effect (see Appendix A for more details): \u03d1 \u03d1x [ h3\u03d1p \u03d1x ] + \u03d1 \u03d1z [ h3\u03d1p \u03d1z ] = \u03d1 \u03d1z [ h3\u03c1g ] + 6\u03bc [ W \u03d1h \u03d1z + U \u03d1h \u03d1x + 2 \u03d1h \u03d1 t ] ", " (17) The balance given in (17) among the pressure fluid flow and the gravitational force, the relative motion (Couette fluid flow) and the, so-called, squeeze flow is set as the basis of an extended parameter variation study in order to allocate both the dynamic and geometric factors, which influence the most the aforementioned dynamic equilibrium. Therefore, a first-order forward finite difference (FD) scheme is utilized (see Appendix B) in order to numerically solve (17) undergoing the following boundary and constraint conditions (see also Appendix A), i.e., p(x, 0) = p(x, l) = p0 and p(0, z) = p(R, z) (18) p(x, z) \u2265 p0 with p0 : atmospheric pressure (19) The two-dimensional representation of the bearing (Fig. 1) should undergo the aforementioned constraints in order for the equivalency regarding the actual three-dimensional representation to hold. Therefore, \u2013 the pressure at both ends of circumferential coordinate (x-coordinate in Fig. 1) should be identical (18), and \u2013 the pressure w.r.t the bearing\u2019s open ends, i.e., z-direction in Fig. 1, should not be less than a predefined pressure, which in most cases suffices to be the atmospheric pressure (19). 3.1 Stochastic-based parameter variation: response functions In regard of the current investigation, the specific multivariate analysis algorithm (MVA) field of stochastic studies [5,13] is selected and implemented along with the numerical solution of the Reynolds equation. In order to sufficiently evaluate the results of each parameter variation study, a set of response functions is selected, the information of which is considered to be rather important w" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000342_j.triboint.2019.105881-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000342_j.triboint.2019.105881-Figure7-1.png", "caption": "Fig. 7. (a) Pad surface temperature measurement point; (b) Pad cross-section for thermocouple installation.", "texts": [ " 2(b)) can be calibrated using the springs below it. The material of the pad surface was a Babbitt metal, which has a high thermal conductivity. The temperature sensor was installed under the pad surface so that the oil film would not be disrupted. Its distance from the pad surface was 4 mm. The \u201cT\u201d type thermocouple was used because its range is \u2212200~+350 \u00b0C, which covered the temperature variation range of the experiment. In addition, it had a small size and was easy to install. The location of the thermocouple on the pad is shown in Fig. 7(a), and the cross-section is shown in Fig. 7(b). The parameter of the tested thrust bearing is shown in Table 3. At the average radius, the circumference of the thrust pad was 420mm and the distance between the two pads was 65mm. A solid example is presented here to demonstrate the influence of the upstream oil film outlet temperature on the downstream oil film. The calculation parameters are shown in Table 4: The calculation model with inlet temperature boundary is shown in Fig. 8. The computational domain includes the oil film and the thrust pads" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001977_j.ijfatigue.2016.09.006-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001977_j.ijfatigue.2016.09.006-Figure6-1.png", "caption": "Fig. 6. Parent material test pieces for tests under oscillatory bending, (mm).", "texts": [ " The results were also taken from the literature [25]. The test pieces were made of hot rolled sheets joined with butt weld, then involved plasma cutting out of test pieces of desired width and finally machining and grinding. In such a manner, test pieces have a geometry identical to that of the first test pieces shown in Fig. 4. These test pieces are example of structural notch. The paper includes also test results [26] related to bending load. In this case, the type of test pieces was parent material test pieces of square cross-section (Fig. 6). The fourth type of test pieces considered are butt joints (Fig. 7) whose results are also taken from the literature [25,27,28]. The weld joints were made of plates t = 10 mm thick (tensioncompression) and t = 30 mm (oscillatory bending). These test pieces are an example of complex notch. Two types of geometric notch are also considered. The first type was subjected to oscillatory bending and prepared as flat elements with geometrical double V-notch. Initially the samples were cutout of rolled sheet 4 mm thick" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001765_1464419315584709-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001765_1464419315584709-Figure3-1.png", "caption": "Figure 3. Position relation of ball and race.", "texts": [ " Dynamic model of ball bearings The dynamic model of ball bearing is developed based on the research of Gupta, which is described briefly in this section, and more details can be found in literature.3,4,23 In present work, some assumptions are employed, as listed below: . Mass and geometric centers are all coincident with each other for each bearing element. . Surface of each bearing element is smooth. Ball\u2013race interactions. The interactions between ball and race are calculated based on their position relation and velocities. The position relation is shown in Figure 3, where victor rr locates the center of the race in inertial frame X,Y,Z\u00f0 \u00de. Victor rr is the center of raceway groove relative to the race center at UNIV OF PITTSBURGH on June 2, 2015pik.sagepub.comDownloaded from described in race fixed frame Xr,Yr,Zr\u00f0 \u00de, and the center of the ball is located by rb in inertial frame. The center of the ball relative to the race center in race fixed frame is rrb r \u00bc Ti r rb rr\u00f0 \u00de \u00f010\u00de where Ti r is the relevant transformation matrix from inertial to race coordinate frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002842_biorob49111.2020.9224374-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002842_biorob49111.2020.9224374-Figure5-1.png", "caption": "Fig. 5. The Dephy ExoBoot ankle exoskeleton used in the physical experiment. A brushless DC motor mounted on a rigid shank assists the user by generating torque through a belt drive transmission (slack in the image) that applies force on a boot-mounted strut. The exoskeleton is securely attached to the user via a shank attachment that transmits the actuator\u2019s torque to the body.", "texts": [ " The natural variability of metabolic rate during walking [15], [30] results in a stimulus noise with a standard deviation about 170% of the initial estimate of JND; binned MOCS would result in a suboptimal sampling plan due to bias with this variability level, whereas SP remains bias-free. Once we had chosen our sampling strategy, we conducted a pilot study using a single subject (male, 20 years old) to validate the methodological choices and obtain the JND of metabolic expenditure. The metabolic rate changes were indirectly imposed by an ankle exoskeleton (ExoBoot, Dephy Inc., Maynard, MA). The exoskeleton (Fig. 5) uses a flat belt drive transmission to apply a plantarflexion torque during walking. Ideally, this torque reduces the biological torque the triceps surae must exert\u2014thereby lowering the total energetic expenditure of the wearer\u2014but it can also cause greater metabolic expenditure by performing negative mechanical work during controlled dorsiflexion. We applied a parametric angle-dependent torque profile that resembled a square pulse to the user via this exoskeleton in which we controlled the onset timing, high-period torque magnitude, and high-period duration of the pulse" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003789_s11694-021-01030-5-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003789_s11694-021-01030-5-Figure5-1.png", "caption": "Fig. 5 a Cyclic voltammograms of GCE-FeCF toward 3.5 \u00b5M tea polyphenols in 0.2\u00a0M PBS with different scan rates. b Plots of peak currents vs. scan rates", "texts": [ " From Fig.\u00a04b, it can be seen that the reduction peak potential of tea polyphenols on GCE-FeCF showed a good linear relationship with pH, and the linear equation was Epc = \u2013 0.0607pH + 0.7120. The slope of 0.060 mV/pH is close to the theoretical value of 0.059 mV/pH, indicating that the reaction of tea polyphenols on GCE-FeCF was a process involving equal electrons and equal protons [43\u201345]. The effect of different scan rates on the electrochemical behavior of 3.5 \u00b5M tea polyphenol solution is shown in Fig.\u00a05. It can be seen that the peak increases gradually when the scan rate is from 20 mV/s to 180 mV/s (Fig.\u00a05a). A linear relationship between the peak current and scan rate can be found in Fig.\u00a05b. The linear equation is: Ipc = 0.00848v + 0.56506 (R2 = 0.9948), which indicates that the electrode process is controlled by the adsorption process [46\u201348]. According to Laviron\u2019s theory [49], the electron transfer number of tea polyphenols reacting on GCE-FeCF can be derived from the following equation: where n is the number of electron transfer, F is the Faraday constant, A is the electrode surface area, \u0413 is the adsorption amount, R is the molar gas constant, T is the thermodynamic temperature, Q = nFA\u0413 is the peak area, v is the scan rate; when the scan rate v = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003021_tec.2020.3045883-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003021_tec.2020.3045883-Figure4-1.png", "caption": "Fig. 4. Geometrical parameters of a cylinder that modeled for the SRM. proposed model block that consists a lumped parameter thermal network (LPTN). According to the charging time the users determine the battery charge current. The winding temperature can be predicted by the proposed model during charging time, and according to the motor insulation class and the rising winding temperature above the allowable temperature, the appropriate decision is made to determine the lower charge current for reducing the winding temperature. With this control algorithm, the battery can be charged in the shortest possible time without deteriorating the winding insulation, this process is illustrated in Fig. 3 by the blue area. In thermal analysis, the geometric symmetry is applicable when the symmetry of initial values and boundary conditions are met in a problem. Therefore, there is a possibility that phase currents or the heat sources are not equal to appoint zero total torque, so one sixteenth (the half of the stator pole) of the 8/6 SRM can\u2019t be used. In this case, a major difference between the thermal model in this paper and those of conventional models is existence of heat transfer between the symmetrical pieces of the SRM. In the following section, an LPTN for one sixteenth of the 8/6 SRM is presented, then the developed model is proposed.", "texts": [ "html for more information. Natural convection between the external surface of the motor and ambient air is defined by Rcov , also between the internal surface and internal air which is described in (9), where h is the heat transfer coefficient and A is area of convection. Rcod demonstrates the heat transfer by conduction between different motor nodes, based on (10) and (11), Rcov,Tan and Rcov,Rad are resistances that represent the heat transfer in tangential and radial directions, respectively [12]. In Fig. 4, geometrical parameters of the cylinder that modeled for the SRM is shown. Rcov = 1 A.h (9) Rcod,Tan = l A.k (10) Rcod,Rad = ln(ro/ri) 2\u03c0kL (11) To obtain a thermal model for the 8/6 SRM when the battery pack is charging, first, the thermal model of one Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 21:21:56 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_22_0002258_s00170-020-04953-3-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002258_s00170-020-04953-3-Figure5-1.png", "caption": "Fig. 5 Grinding movements of the machine", "texts": [ " In the grinding process, the B axis and point o3 is actual rotation center of the workbench, while point o2 is virtual rotation center of the workbench. The actual grinding center distance is a' and the theory center distance expresses by a. In Fig. 4, it can be known that the grinding center distance a' is much smaller than the theory center distance a. Therefore, the machine structure with this virtual rotation center grinding principle is much smaller than the actual rotation center grinding principle. Based on the virtual rotation center grinding principle the four-linkage grinding movements are shown in Fig. 5. Here, point P1 is the actual rotation center of the workbench at the beginning of grinding, while point P2 represents the actual rotation center of the workbench at the ending of grinding. Point P1 shows value of (xp1, zp1) in the machine coordinate system, and point P2 contains value of (xp2, zp2) in the machine coordinate system. \u03b812 expresses the angular displacement of B axis, when \u03c612 indicates angular displacement of generating plane on the main basic circle. According to the Eqs. (10) and (11), as well as the Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001142_sii.2014.7028137-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001142_sii.2014.7028137-Figure3-1.png", "caption": "Fig. 3. Control box of ERIK", "texts": [ " It is thus possible to realize real-time resistance control. ER fluids have been used in various appplications, including dampers, shock absorbers, and clutches in addition to controlled brakes [3], [4], [5], [6]. The appearance of ERIK is shown in Fig. 2. ERIK comprises a control box and slider part. The slider part comprises a timing belt and roller platform, while the control box houses a motor, ER fluid brake, and encoder pulley, and is connected to a laptop computer. The interior of the control box is shown in Fig. 3. The timing belt of the slider part is connected to pulley 1. Pulley 978-1-4799-6944-9/14/$31.00 \u00a92014 IEEE 779 2 is connected to the ER fluid brake and motor. In addition, a pin is inserted in the ER fluid brake, and the ER fluid brake acts as a brake when the pin is inserted and as a clutch when the pin is removed. An example of the use of ERIK in strength training is shown in Fig. 4. Here ERIK provides a resistance that is greater than its own weight. As discussed above, the control of ERIK can be switched between a clutch mode and brake mode" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001083_1.4027926-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001083_1.4027926-Figure2-1.png", "caption": "Fig. 2 Aero-engine rotor experimental rig", "texts": [ " In comparison to a real aero-engine, the tester has the following features: (1) its size is one third that of a real aero-engine, and its shape is similar to the stator of an aero-engine; (2) its internal structure is simplified: the core machine is simplified to a 0-2-0 support structure, which is a rotor with two disks supported on two bearings, where the outsides of the two bearings have no other disks, and the support stiffness is adjustable to adjust the dynamic characteristics of the system; the multistage compressor is simplified to a singlestage disk structure; the compressor blade is simplified to an inclined plane shape; (3) its shaft is solid and rigid and its maximal rotating speed is 7000 rpm; and (4) the rotor is driven by the motor, and the flame flask is canceled. Therefore, the aero-engine rotor tester is a single-rotor system model. A full-scale photo is shown in Fig. 2(a), and a section drawing is shown in Fig. 2(b). In the aero-engine rotor experimental rig, the compressor part of the rotor is supported by a roller bearing, and the turbine part of the rotor is supported by a ball bearing. In this study, faults are produced on the ball bearings, and fault simulation experiments are carried out. To study the difference between the responses of the bearing house and casing caused by ball bearing faults, and analyze the sensitivity of the casing response to the ball bearing faults, impulse response experiments are first carried out" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001651_s00170-016-8545-0-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001651_s00170-016-8545-0-Figure3-1.png", "caption": "Fig. 3 Blisk EBM model design", "texts": [ " Regardless of the EDM machining error, the EBM forming error and the machining reference establishment error during the EDM eroding process are the twomain factors which influence the blade accuracy most in EBM/EDM combined manufacturing. The most direct way to get the required blade profile accuracy is to leave very large machining allowance on the EBM blank. Nevertheless, the combined manufacturing aims at reducing the machining allowance for EDM to save time. It is not worth the candle if the machining allowance is too much. The reasonable machining allowance should be a little larger than the EBM forming error. For a shrouded blisk as shown in Fig. 3, three machining references need to be established for EDM eroding. They are the axial reference, the radial one, and the circumferential one. An axial end face and a cylindrical face on the shrouded blisk are finish machined by cutting and are selected as the axial and radial machining references for the following EDM process, respectively. EBM model modification, such as extending the end face into large plane and unifying the diameter of cylinder, is always conducted to get better EDM forming results", " Visual identification needs a few photos of the blisk blank after it is put on the EDM machine tool. These photos are taken in the same place and while blisk is rotating to keep the angles between blades and electrode are different. A transformation can be built by recognizing the changes from these objects to their images based on the photos. Then the relationship between the blades and the electrode can be recognized from these photos. Therefore, visual identification is very convenient to extract features from the near-net formed blisk. Figure 3 shows the design of the blisk EBM model. The theoretical EDM machining allowance of every single surface on blisk is set as 0.5 mm, which is a little larger than the EBM machining error 0.3 mm. This ensures the EBM blisk blank envelops the theoretical profile of the final blisk even if the maximum error of EBM forming happens. Figure 4 presents the blisk blank formed by EBM. Visual identification is conducted when blisk cutting is finished. At that time, the axial and radial machining references of the blisk have been machined" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure11-1.png", "caption": "Fig. 11. Differential displacement of a screw a\u0302 around the screw axis b\u0302 with the differential angle dv.", "texts": [ " A zero-pitch unit screw (dual unit vector) describes an oriented line in space and is in motor notation given by u\u0302 \u2261 [ u ue ] = [ u r\u0303u ] , |u| = 1, r \u00d7 u \u2261 r\u0303u (68) with the unit vector of direction u and its moment ue with respect to the reference point O0, see Fig. 10a. The zero-pitch unit screw (68) fulfils the PL\u00dcCKER condition uT ue = 0. (69) A general screw comprises an oriented line in space with an associated pitch h (Fig. 10b), a\u0302 \u2261 [ a ae ] = [ a r\u0303a + ha ] = [ a ae\u22a5 + ae\u2016 ] , |a| = a. (70) Differential displacement of a screw In the following a differential displacement of a screw a\u0302 around the screw axis b\u0302 with the differential angle dv is considered, see also Fig. 11, with a\u0302 = [ a r\u0303aa + haa ] , b\u0302 = [ b r\u0303bb + hbb ] , |a| = |b| = 1. (71) The differential displacement of the screw a\u0302, da\u0302 = [ da d\u0303raa + r\u0303ada + hada ] \u2261 [ da dae ] , (72) is composed of the differential rotation of the vector a around the rotation axis b with the differential angle dv da = b\u0303adv (73) and the differential increment of the dual part dae of a\u0302. The differential displacement dra of the axis of a\u0302 in Eq. (72) can be composed of the rotational part drba\u22a5 and the translational part drba\u2016 of the vector rba = ra \u2212 rb, thus dra = b\u0303(ra \u2212 rb)dv\ufe38 \ufe37\ufe37 \ufe38 drba\u22a5 + hbbdv\ufe38 \ufe37\ufe37 \ufe38 drba\u2016 = (\u0303bra + r\u0303bb + hbb\ufe38 \ufe37\ufe37 \ufe38 be )dv" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure3.18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure3.18-1.png", "caption": "Fig. 3.18 Visual comparison of the scan quality of the SICK LMS and Sick IVP Ruler. a The scan of the LMS has less occlusions but a much higher amount of noise. b The Ruler has much less noise but a significant amount of holes caused by shadows between laser line and camera. c 5 located piston rods in the LMS scan, showing the robustness of the approach against noise. d 5 located piston rods in the Ruler scan, showing the robustness of the approach against occlusions", "texts": [ "14 Plot of the results of the RANSAM algorithm applied to isolated objects with different noise levels . . . . . . . . . . . . 31 Figure 3.15 Set of virtually scanned test scenes . . . . . . . . . . . . . . . . . . 32 Figure 3.16 Plot of the results of the RANSAM algorithm applied to the random viewpoint scenes . . . . . . . . . . . . . . . . . . . . . 33 Figure 3.17 Three vision sensors mounted on the linear axis in the robot work cell . . . . . . . . . . . . . . . . . . . . . . . . . . . . 34 Figure 3.18 Visual comparison of the scan quality of the SICK LMS and Sick IVP Ruler . . . . . . . . . . . . . . . . 35 Figure 3.19 Effect of the edge based localization attempt when locating critical planar objects . . . . . . . . . . . . . . . . . . 36 Figure 4.1 Gripper kernel KG used for gripper pose hypotheses generation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41 Figure 4.2 Pick pose estimation by correlation . . . . . . . . . . . . . . . . . . 42 Figure 4.3 Estimation of the gripper orientation \u03a6", " The second important difference is that the LMS400, as it is not triangulation based and therefore does not need a big baseline, produces significantly less shadows at the cost of a lower accuracy. The data sheet gives a systematic error of \u00b14mm and a statistical error of \u00b110mm, which is comparable to the highest simulated amount of noise. Whereas the Ruler is specified with a typical height resolution of 0.4mm, 34 3 3D Point Cloud Based Pose Estimation which is comparable to the noise free simulation. A visual comparison of the details captured by each of the scanners can be seen in Fig. 3.18. The LMS has an operating range of 0.7 . . . 3m, the Ruler has an operating range of 0.28 . . . 1.28m, i.e. both are suited for the experimental setup as well as for a possible industrial setup. The first set of experiments was performed using the LMS with its high amount of noise. The second set of experiments was performed with the Ruler, which is more accurate. As objects served a set of industrial metal parts, which were scrambled unmixed in the bin. The task was to locate an object, to find a secure grasp pose and to place the object on a seating" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002697_01691864.2020.1810772-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002697_01691864.2020.1810772-Figure1-1.png", "caption": "Figure 1. A serial robot.", "texts": [ " Moreover, the method of updating parameters is also explicit and unified. This paper is organized as follows. Section 2 presents the modified Adjoint error model, then the complete calibration process is shown in Section 3, and all the details are given for the calibration model. For verifying the effectiveness of the calibration method, the simulations are conducted in Section 4. A brief conclusion is drawn in Section 5. Finally, the mathematical background not mentioned in the main body is provided in the \u2018Appendices\u2019. As shown in Figure 1, there is an n degrees of freedom (DOF) serial robot, with revolute and prismatic joints, and we fix the coordinate frame of the S and T on the base and end-effector, respectively. According to the POE model, the pose of the frame T with respect to the frame S can be described as [12] fst(q) = e\u03be\u03021q1 \u00b7 \u00b7 \u00b7 e\u03be\u0302iqi \u00b7 \u00b7 \u00b7 e\u03be\u0302nqngnst0 , (1) where the joint twist \u03be\u0302i \u2208 se(3) is associated with the joint i, and qi \u2208 is the joint variable, and gnst0 \u2208 SE(3) is the initial pose of the frame T in terms of the frame S, and the exponential of \u03be\u0302iqi can refer to Appendix 1 for details" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001820_0954408916652648-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001820_0954408916652648-Figure1-1.png", "caption": "Figure 1. The contact region of a single asperity.", "texts": [ " Majumdar and Tien8 presented a fractal representation method on the rough surface by using the W\u2013M function, which can define a roughness surface with fractal parameters. Majumdar and Bhushan9 presented the famous M\u2013B fractal contact model to calculate the total real contact area and normal load based on the two-dimensional W\u2013M function and Hertz theory. For the MB model, the contact of two rough surfaces is assumed as a contact of a rigid surface and a rough surface. The rough surface is composed by many asperities with different sizes. Figure 1 shows the contact region of a single asperity. The asperity from rough surface is squeezed against the rigid plane. For the purpose of further improving the prediction accuracy, some researchers tried to modify the classical M\u2013B fractal model. Wang and Komvopoulos10,11 presented an improved size-distribution function of the micro-contact spot with a domain extension factor , and divided the contact deformation of the asperity into three stages (the totally elastic stage, the elastic-plastic stage and the totally plastic stage)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002433_j.ijsolstr.2020.03.020-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002433_j.ijsolstr.2020.03.020-Figure4-1.png", "caption": "Fig. 4. The elementary module of a repetitive frame with overlapping.", "texts": [ " In structural engineering, repetitive frames at a wide range of cales may be modelled by the repetition of modules composed y few rigid rods connected by rotational joints \u2013 see for example Milton, 2013 , Guest and Hutchinson, 2003 , Tanaka and Shibutani, 009 , Tanaka et al., 2012 ). In addition, one of the most relevant eatures of repetitive frames, such as those used to represent many d i i i i s d w e t w o \u03b8 C e e g i m V H I f l types of tailor-made metamaterials, is that we can adjust some mechanical properties in order to obtain a pretended constitutive behaviour ( Bertoldi et al., 2017 , Findeisen et al., 2017 ). Within this context, in the present section the rods system presented in Fig. 4 is analysed in the buckling and post-buckling ranges. In this system, the secondary bars [AB] and [CD] overlap the main ones [BC] and [DE], and we suppose henceforth that AB + CD is always smaller than BC or DE . The relevant mechanical properties \u2013 rods lengths and joints stiffness coefficients \u2013 are varied in order to study their influence on the global tensile buckling behaviour. Alternative geometries for the module could have been adopted, which we leave for future works, but this specific geometry was chosen in order to allow its sequential repetition with no changes at all in each prolongation step: for the module depicted in Fig. 4 - a), the repetition is simply made by attaching point A of the subsequent module to point E of the previous one. In the following, we are primarily interested in how overlapping generates tensile buckling for the repetitive system made up of the sequential reproduction of the module of Fig. 4 , with a special focus on the original behaviours that may arise, such as compound or sequential buckling, and on the smaller sensitiveness to the number of modules when compared to compressive buckling. To this end, several examples are presented in the following, in order to illustrate the original buckling behaviours that may arise from overlapping, together with the influence of the module\u2019s repetition and of the mechanical properties on the tensile buckling and post-buckling phenomena. 3.2. The analysis of the single module 3.2.1. The energy formulation and general stability analysis The effect of the module\u2019s repetition on the tensile buckling behaviour of the global system can only be properly assessed if at first the behaviour of the single module is thoroughly analysed. To this end, let\u2019s consider the module depicted in Fig. 4 , made up by four rigid links connected by elastic rotational springs at points A, B, C, D and E, each one with stiffness coefficients K i , i = A, B , \u2026, E . Point A remains fixed as the frame deforms under the action of an axial force at point E and the module is cinematically escribed by: ) PointB : { w B = ( \u22121 ) 1 AB sin \u03b8AB u B = ( \u22121 ) 1 AB ( cos \u03b8AB \u2212 1 ) (22) i) PointC : { w C = w B + ( \u22121 ) 2 BC sin \u03b8BC u C = u B + ( \u22121 ) 2 BC ( cos \u03b8BC \u2212 1 ) (23) ii) PointD : { w D = w C + ( \u22121 ) 3 CD sin \u03b8CD u D = u C + ( \u22121 ) 3 CD ( cos \u03b8CD \u2212 1 ) (24) v) PointE : { w E = w D + ( \u22121 ) 4 DE sin \u03b8DE = 0 u E = u D + ( \u22121 ) 4 DE ( cos \u03b8DE \u2212 1 ) (25) Crucial at this stage is the repetitive character of the structural ystem and of its mathematical modelling", " onsequently, only three of these four degrees of freedom are linarly independent, and the module\u2019s total potential energy considrs this kinematic restraint of expression ( 25 ) by means of the Larange Multipliers Method ( Ikeda and Murota, 2010 ) \u2013 the resultng energy function is as follows, where \u03bb denotes the Lagrange ultiplier: = K A 2 \u03b82 AB + K B 2 ( \u03b8AB \u2212 \u03b8BC ) 2 + K C 2 ( \u03b8CD \u2212 \u03b8BC ) 2 + K D 2 ( \u03b8CD \u2212 \u03b8DE ) 2 + K E 2 \u03b82 DE \u2212 P \u00b7 u E + \u03bb w E (26) The application of the stability procedures ( Thompson and unt, 1973 , Thompson and Hunt, 1984 , Godoy, 20 0 0 , Hunt, 1981 , keda and Murota, 2010 , Sim\u00e3o et al., 2012 ) to the potential energy unction ( 26 ) leads to the following generalized eigenvalue probem for the evaluation of the critical loads along the fundamental Table 1 Adopted properties for the elementary frame module of Fig. 4 . Case Lengths Case Stiffness coefficients AB BC CD DE L tot K A \u2261 K 0 K B \u2261 K 1 K C \u2261 K 2 K D \u2261 K 3 K E \u2261 K 4 A 1 3 1 3 4 1 1 1 1 1 1 B 1 2 1 2 2 2 1 0.1 1 0.1 1 C 1 4 1 4 6 3 1 1 0.1 0.1 1 D 1 4 1 2 4 4 0 1 1 1 0 5 0 1 0 1 0 p H w a H H s d f m e b d t t 3 b p t o u \u03ba s o a a a 3 j s e t p p a T j i n l F \u03ba ath: F P \u00b7 \u23a7 \u23aa \u23aa \u23a8 \u23aa \u23aa \u23a9 \u03b8AB \u03b8BC \u03b8CD \u03b8DE \u03bb \u23ab \u23aa \u23aa \u23ac \u23aa \u23aa \u23ad = ( H 0 \u2212 P \u00b7 H 1 ) F P \u00b7 \u23a7 \u23aa \u23aa \u23a8 \u23aa \u23aa \u23a9 \u03b8AB \u03b8BC \u03b8CD \u03b8DE \u03bb \u23ab \u23aa \u23aa \u23ac \u23aa \u23aa \u23ad = \u2212\u2192 0 (27) here the parts H 0 and H 1 of the energy\u2019s Hessian matrix H FP are s follows: 0 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a3 K A + K B \u2212K B 0 0 \u2212AB \u2212K B K B + K C \u2212K C 0 BC 0 \u2212K C K C + K D \u2212K D \u2212CD 0 0 \u2212K D K D + K E DE \u2212AB BC \u2212CD DE 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a6 (28) 1 = \u23a1 \u23a2 \u23a2 \u23a2 \u23a3 AB 0 0 0 0 0 \u2212BC 0 0 0 0 0 CD 0 0 0 0 0 \u2212DE 0 0 0 0 0 0 \u23a4 \u23a5 \u23a5 \u23a5 \u23a6 (29) The relevant critical states along the fundamental path correpond to the eigenvalues of the eigenproblem ( 27 ) and, due to its imension, an explicit formula for the lowest tensile critical load or a generic case is no longer feasible", ", 2012 ) must be applied to ach example, to compute its tensile buckling and post-buckling ehaviours. In the following, several examples are presented in orer to illustrate the potential of overlapping in creating original ensile instability-based behaviours of bifurcational type in repetiive frames. t ig. 5. The elementary module of a repetitive frame with overlapping: variation of the co ref and with the lengths amplitude factor ref . .2.2. Example 1: comparison between tensile and compressive uckling behaviours The first example considers a single module of Fig. 4 whose roperties follow case A1 in Table 1 . Fig. 5 depicts the lowest posiive (tensile) and the lowest negative (compressive) buckling loads, btained by solving the eigenproblem ( 27 ) for a the single modle, considering that the rods\u2019 lengths are all multiplied by factor ref and the joints stiffness coefficients are all multiplied by factor ref .. It is thus found that the two lowest critical loads have oppoite signs: one is negative, corresponding to compression, and the ther is positive, hence related to tensile buckling", " 14 , hich presents the frame\u2019s post-buckling shapes, it becomes clear hat the first critical state is related to the instability of bars [CD] nd [DE], while the second critical state refers to the buckling of he remaining bars of the module \u2013 the shorter bars buckle first hen the module is tensioned. .3. Overlapping as a means to create repetitive systems with egative thermal expansion coefficient Materials with negative thermal expansion coefficient have een deeply studied in recent years ( Wei et al., 2018 , Li et al., 019 ), owing to a wide range of promising applications in the enineering practice. In the structural systems composed by the seuential repetition of the module depicted in Fig. 4 this property an be achieved quite easily. Let\u2019s imagine this module is comosed by two distinct materials \u2013 material a for bars [AB] and [CD] ith positive thermal expansion coefficient \u03b1a , and material b for ars [BC] and [DE] with positive thermal expansion coefficient \u03b1b . hen the whole module is exposed to a uniform variation of temerature \u03b8 , considered without loss of generality to be positive, oint E moves along the x -axis of the following amount: E = [ \u2212\u03b1a ( L AB + L CD ) + \u03b1b ( L BC + L DE ) ] \u03b8 (30) This displacement is negative, implying the shrinkage of the odule as a whole with heating, when: a ( L AB + L CD ) > \u03b1b ( L BC + L DE ) (31) hich leads to a module with negative global thermal expansion oefficient", " This finding clearly suggests overlapping also as a route o generate frames with negative thermal expansion. 3 3 f t e t t s p o .4. The repetition of the module .4.1. The energy formulation and the Hessian matrix along the undamental path In nature and in structural engineering repetitive frames are ofen adopted, owing to their efficiency to carry load and to their ase of erection ( Guest and Hutchinson, 2003 , Tanaka and Shibuani, 2009 , Tanaka et al., 2012 , Gibson and Ashby, 1997 ). By repetiion of the module depicted in Fig. 4 we consider merely to attach equentially node A of the subsequent module to node E of the revious one, thus making both edge nodes to coincide. On the ther hand, looking at expressions ( 22 ) to ( 25 ) it is immediately b \u23a7\u23aa\u23a8 \u23aa\u23a9 w i o i f noticed that the kinematic description of such a module has a recursive character: for a generic node of the frame, any displacement is obtained by adding a quantity to the correspondent displacement of the anterior node in the sequence, and this quantity follows a constant pattern along the module. This recursive character applies also to the module\u2019s repetition and represents mathematically the repetitiveness. For the sake of simplicity, it is henceforth advantageous to replace the node\u2019s reference letter by a number related to the node\u2019s ordering along the sequence: for the module depicted in Fig. 4 , node A becomes node 0, node B is node 1, node C is node 2 and so on. Furthermore, the kinematic variable \u03b8 and its corresponding bar are related to the index number of the bar\u2019s end node, so that \u03b8AB = \u03b81 , \u03b8BC = \u03b82 , and so on. So, for a generic node k of the repetitive structure its displacement is given y: u k = k \u2211 i =1 ( \u22121 ) i L i ( cos \u03b8i \u2212 1 ) w k = k \u2211 i =1 ( \u22121 ) i L i sin \u03b8i (32) here L i represents the i th bar\u2019s length, whose end node is node , and k varies between 1 and n N , where n N is the total number f bars, given by n N = 4 n M ", " This property is of major relevance and clearly distinuishes the compressive buckling behaviour from the tensile one. s f s D n a t a p u u s It arises from the fact that the compressive critical load is associated with the global instability of the frame as a whole, hence it depends on the frame\u2019s total length that increases when additional modules are added to the system, while the tensile one is related mainly to the overlapping length that is kept constant as the number of modules is increased. 3.4.3. Example 5: general tensile post-buckling analysis Let\u2019s consider a frame with two equal modules of Fig. 4 , both following the properties listed in Table 1 . While keeping the numerical problem within a reasonable dimension, this structural ystem already shows some features that are observed in longer rames due to the larger repetition of the modules. Fig. 16 decribes the postbuckling behaviours, for lengths\u2019 cases A, B, C and and for stiffness coefficients\u2019 cases 1 to 3, by plotting the end ode\u2019s axial displacement u , divided by the frame\u2019s total length, gainst the applied load P . It becomes evident that the manipulaion of the bars lengths and of the springs stiffness coefficients is route to tailor-made original elastic constitutive relations in the ost-buckling range", " The properties affected by this global variation may be of geometric type (such as dimensions and shape) and/or of mechanical type (such as the material\u2019s rigidity or strength), and this strategy is used to adapt the frame to local factors, such as peaks of stresses, or to adapt the frame\u2019s shape to fulfil any existing design requirement. Furthermore, global variations of the mechanical properties are often used to represent for example imperfections in cellular structures and materials ( Silva and Gibson, 1997 , Zhu et al., 2001 ), but for the moment we pass over these practical applications and turn instead our attention to the effects these variations of the structure\u2019s mechanical properties provoke on the tensile buckling of the repetitive frame made up by the sequential repetition of the module depicted in Fig. 4 , namely the influence of varying the joints stiffness coefficients and the rods lengths. The examples showed in section 3 , all related to perfect repetition, are now taken as reference cases to which the global and smooth variations of rods lengths and joints stiffness coefficients are applied. So, it is assumed that the application of a smooth variation to a specific mechanical property consists merely in multiplying the property\u2019s original value for element i of the reference case by a global variation factor k var,i , in the present case ssumed to be given by: v ar,i = 1 + \u03baamp sin ( i \u2212 1 n re f \u2212 1 \u03c0 ) , i = 1 , ", " We leave for future developments a wider parametric study n the influence of these variation parameters, in amplitude and in hape, on the global frame\u2019s tensile buckling behaviour. To sum up nd for the cases listed in Table 1 , element i to be affected by exression ( 40 ) can be either a bar, so expression ( 40 ) multiplies the th bar\u2019s initial length, or a rotational spring, and expression ( 40 ) ultiplies the i th joint\u2019s stiffness coefficient. In the following, the onsequences of such variations on the tensile buckling behaviour f the frame built up by the sequential repetition of the module epicted in Fig. 4 are assessed by the analysis of relevant examles, and some original behaviours will be found, completely disinct from the ones observed in the correspondent reference cases. hese findings clearly establish the global variation of mechanical roperties as a strategy to create original constitutive behaviours n instability-based repetitive frames. .2. Illustrative examples .2.1. Example 7: variation of the rods lengths and the joints stiffness oefficients in a 7 modules frame following Case 1 Let\u2019s consider a frame built up by the 7 times repetition of the odule presented in Fig. 4 , whose properties follow case A1 of Table 5 The tensile critical loads for the 7 modules system for all variation cases (boxed cases L0-S0 are the reference cases). Case A1 Case A3 L0 L1 L2 L0 L1 L2 S0 1.345 1.561 1.099 0.161 0.184 0.132 S1 1.039 1.342 0.840 0.124 0.161 0.100 S2 1.558 1.682 1.347 0.185 0.193 0.161 T t T c c e o t t c a f r s c l e t r a d t p r 4 c s i E t s fl w C 2 o j t t a c s c t d 4 s e o a w w f t i r able 1 . The lengths and joints stiffness coefficients affected by he variations coefficients given by expression ( 40 ) are listed in ables 3 and 4 for all cases, respectively, and the correspondent ritical loads are presented in Table 5 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003079_0954407020947494-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003079_0954407020947494-Figure2-1.png", "caption": "Figure 2. Phase diagram of the motor speed and the wheel hub speed near the original system subcritical Hopf bifurcation point (a2 = \u20133.300) under different initial values: (a) initial values are [0.1 0.1 0.1 0.1] and (b) initial values are [2.5 2.5 2.5 2.5].", "texts": [ " Similarly, the result of equation (17) is c2 = 0:0006894 when the bifurcation parameter is a 22 = 3:231, which means that the bifurcation type occurring at this point is subcritical Hopf bifurcation. Above all, with the influence of the primary term friction coefficient (a2) of the system sliding friction moment considered, the torsional vibration stability zone of the wheel-side direct-driven transmission system is determined as [\u201318.097, \u20133.231]. When the bifurcation parameter is taken as a2 = 3:300 which is close to subcritical Hopf bifurcation point from the stable region, it is seen from Figure 2 that the stability of the system torsional vibration responses is determined by the system initial values. When the initial values of the system state variables are [0.1 0.1 0.1 0.1] which is close to the equilibrium point, the torsional vibration responses of the wheel-side direct-driven transmission system are asymptotically stable, which is shown in Figure 2(a). However, when the initial values of the system state variables are [2.5 2.5 2.5 2.5] which is relatively far from the equilibrium point, the system torsional vibration responses are unstable and the system phase diagram is gradually divergent, which is shown in Figure 2(b). For the subcritical bifurcation point, the system torsional vibration responses will change from stability to instability suddenly, which is harmful to the safe operation of the wheel-side direct-driven transmission system. Therefore, it is necessary to avoid the occurrence of the subcritical Hopf bifurcation phenomenon. For the supercritical Hopf bifurcation point, the torsional vibration responses of the system are stable, so only the responses amplitudes need to be controlled effectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000623_0954406219893721-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000623_0954406219893721-Figure5-1.png", "caption": "Figure 5. Completive refining of the mesh density on the loaded side of the tooth surfaces: (a) involute internal gear and (b) novel HCR internal gear.", "texts": [ " A precise geometric model is a good basis for FE simulation. What\u2019s more, a relatively fine mesh density is also critical to simulate the non-linear contact deformation faster and better between the tooth surfaces. Kiekbusch et al.12 introduced two methods to refine the mesh in the area of contact during the simulation process. One is to adaptively refine the mesh at the positions where the tooth surfaces contact occurs. Another is to refine the complete tooth surfaces on the loaded side of the tooth. Figure 5 shows the mesh refinement by using plane stress elements CPS4R38 for the involute internal gear and the novel HCR internal gear with the latter method. In the gear FE model, a surface-to-surface contact property is established between the meshing tooth surfaces, and the two contact faces are defined using the master\u2013slave configuration. Wherein, the pinion contact surface is disposed as a main surface, and the internal gear contact surface is disposed as a slave surface, and a finite slip algorithm is used", " Constrain all degrees of freedom of the center reference point of the pinion, and apply a small input torque (for example, 10Nm) to the internal gear to establish the contact relationship at the meshing tooth surfaces while avoiding drastic changes in the contact state. 2. Change the constraint state of the center reference point of pinion and apply an angular displacement constraint at the center of its rotation within the time specified in the analysis steps. For the involute internal gear pair, make the pinion turn over 0.9 radians (about six teeth); for the HCR internal gear pair, make the pinion turn over 1.6 radians (about 10 teeth), as shown in Figure 5. 3. Change the small load of the above mentioned and then apply the real large loads such as 300Nm, 400Nm, and 500Nm under different working conditions. 4. The normal contact force of the gear pairs and the angular displacement of pinion and internal gear around their respective rotational centers are extracted. Comparison on the torsional mesh stiffness and contact ratio between involute gear and HCR gear Influence of input torque on torsional mesh stiffness It is difficult to accurately distinguish and extract each elastic deformation component from the result of FE models" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003387_icccr49711.2021.9349369-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003387_icccr49711.2021.9349369-Figure2-1.png", "caption": "Figure 2. The model of the ballbot: (a) Top view of the ballbot; (b) Ballbot is approximated into two decoupled plant.", "texts": [ " Additionally, the following assumptions are made [2]: \u2022 Rigid Body Assumption: In this model, all the objects are considered as rigid bodies, which means all the deformation should be neglected. \u2022 Friction Assumption: Friction between the wheel and the floor and between the wheel and the body is modeled as pure viscous damping. Forces due to static friction and nonlinear dynamic friction are neglected. \u2022 No Slip Assumption: There is no slip between the ball and floor and between the ball and the wheel. The three contacting points on the ball form a supporting triangle, which is essential to the system. Consider two cases of the ballbot shown in Fig. 2 (a). As the first case, the ballbot tilt a small angle and its center of gravity is directly above the supporting triangle, which guarantees that all wheels are contacting with the ball. However, for some reason, the inclination angles of a ballbot increase to a large extent, leading to the center of gravity out of the supporting triangle. With the gravity torque, the contact condition and the friction change a lot, thus the assumptions we made when modeling the robot are no longer satisfied. More specifically, there is a practical limit on the magnitude of the tilt angle", " Generally speaking, the body has 2 DOFs: the inclination angle and the direction to which the robot 1) Simplified mathematical model: A nonlinear 3D model of the system dynamics of ballbot has been analytically derived in [2], where the wheel dynamics are neglected. In this model, the three-dimensional system is approximated in three 3 Authorized licensed use limited to: California State University Fresno. Downloaded on June 25,2021 at 20:42:47 UTC from IEEE Xplore. Restrictions apply. decoupled planes. As shown in Fig. 2 (b), the median sagital plane (xz-plane) and the median coronal plane (yz-plane) related to the system moving forward or backward on the flat floor are identical and share same equations of motion. Wheels are simplified as one actuating wheel that fifits in a plane. The third plane (xy-plane) describes the rotation around the z-axis in the body fixed reference frame, to which we pay less attention in balance control. Under these circumstances, the controller for the three-dimensional system can be designed by analysing and designing independent controllers for the two separate and identical systems" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003600_09544070211019227-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003600_09544070211019227-Figure2-1.png", "caption": "Figure 2. Rotating Timoshenko beam generalized coordinates.", "texts": [ " The matrix G is the gyroscopic matrix, and KO is the stiffness matrix of centripetal acceleration resulting from carrier rotation. Kb is the bearing stiffness matrix. Km is the gear meshing stiffness, which is composed of a mean value and a time-varying value. The matrices G, KO, Kb and Km for SSPGs can be given in the appendix of Cooley and Parker.14 Modelling of the rotational Timoshenko beam Intermediate shafts in TSPG systems are usually solidor hollow-shaft rotors. Consider a beam element with generalized coordinates, as shown in Figure 2. Here, vz is the axial displacement, vx and vy are the lateral displacements and ux, uy and uz are the rotations with respect to the x-, y- and z-axes, respectively. The Timoshenko beam incorporates shear forces and rotary inertia.32,33 Based on the generalized coordinates given in Figure 2, the equations of motion for the rotational Timoshenko beam can be written as rA \u22022vx \u2202t2 + kAG \u2202uy \u2202z \u22022vx \u2202z2 = fx(z, t) \u00f03\u00de rA \u22022vy \u2202t2 kAG \u2202ux \u2202z + \u22022vy \u2202z2 = fy(z, t) \u00f04\u00de rI \u22022ux \u2202t2 +2ObIr \u2202uy \u2202t EI \u22022ux \u2202z2 + kAG ux + \u2202vy \u2202z =0 \u00f05\u00de rI \u22022uy \u2202t2 +2ObIr \u2202ux \u2202t EI \u22022uy \u2202z2 + kAG uy + \u2202vx \u2202z =0 \u00f06\u00de where r is the density, E is Young\u2019s modulus, G is the shear modulus, A is the cross-sectional area, I is the cross-sectional area moment of inertia, k is the shear coefficient and Ob is the rotational speed of the beam" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001208_gt2014-26176-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001208_gt2014-26176-Figure8-1.png", "caption": "Figure 8 Contact of interference model", "texts": [ " But when the displacement increases enough, the contact area will keep stable, about half the area of the cylindrical surface. As a result, the contact stiffness will increase little with the displacement u and become stable. In the condition of 0 clearance contact, the contact area keeps in the half part of the cylindrical surface, which almost does not change with the displacement u. Thus, the contact stiffness changes little with u. In the condition of penetration contact, when u is less than - R , the area of contact surface keeps constant, because of the initial penetration, as shown in Figure 8. Thus, the contact stiffness changes a little with u. When the two surfaces first separate at the upside of the cylindrical surface, the contact area decreases rapidly with u. As a result, the contact stiffness will decrease rapidly with u. But when the displacement increases enough, the contact area will keep stable, about half the area of the cylindrical surface. As a result, the contact stiffness will decrease slowly with u and become stable. From the analysis above, it is known that the contact stiffness keeps stable in the condition of large loading" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000530_jzus.a1900163-Figure23-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000530_jzus.a1900163-Figure23-1.png", "caption": "Fig. 23 Experimental device of mechanical seals", "texts": [ " It shows that the minimum film thickness increases first and then decreases with the increase of dimple depth. When 2 \u03bcm\u2264hp\u22645 \u03bcm, the minimum film thickness is larger than 3\u03c3, so that the textured surfaces are in full film lubrication. This means the hydrodynamic effect of the dimples is stronger and the load-carrying capacity is larger when 2 \u03bcm\u2264hp\u22645 \u03bcm. The maximum radial clearance occurs when hp is equal to 4 \u03bcm; it is also the optimal dimple depth for the strongest hydrodynamic effect. The experimental device is shown in Fig. 23. The seal chamber contains double sets of the same mechanical seal, and the upper one is the test seal. There are seven holes on the back of the stator, which are used to measure the temperature of the sealing surface. The diameter of the holes is 2 mm, and the distance between the holes and the sealing surface is 0.5 mm. The surface of the stator with dimples is Yang et al. / J Zhejiang Univ-Sci A (Appl Phys & Eng) 2019 20(11):864-881 878 shown in Fig. 24, the depth of dimples is about 5 \u03bcm, and the dimples are produced by laser surface texturing (LST) technology" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003425_j.ins.2021.02.037-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003425_j.ins.2021.02.037-Figure1-1.png", "caption": "Fig. 1. A schematic diagram of vehicle path planning.", "texts": [ " More information about geometric continuity can be found in [1,2]. In this section, some preliminaries will be introduced. In Section 2.1, a method of dividing curve path is proposed in consideration of curvature. Section 2.2 presents a specified velocity formula which can avoid sudden change of acceleration while ensuring continuous velocity. Essentially, motion planning can be viewed as an autonomous vehicle to trace the assigned path from its current configuration to the determined one (shown in Fig. 1). Suppose that there exists an assigned path whose arc length can be denoted as s. Its curvature q s\u00f0 \u00de can be considered as a function of s. Meanwhile, sf denotes the total length of the path, and tf is the total travel time of vehicle. In general, the assigned path is generated by path planning module of autonomous vehicle, so its curvature can be seen as the prior information during the phase of velocity planning. As previous studies have shown in [4,5], the curvature of path would limit the maximummoving velocity of vehicle if the maximum normal acceleration is considered", " (11) means that the maximum (minimum) acceleration of vehicle is a fixed value after the initial and final velocities on the assigned path are determined from the specified velocity formulas shown in Eqs. (2)\u2013(4). This section introduces the problem formulation and a modified PPSO algorithm is proposed to obtain the final velocity profile. In Section 3.1, the premier problem is transformed into an optimization problem. And in Section 3.2, a modified PPSO algorithm is proposed to solve the optimization problem with high dimension variables while the neighbor variables are subject to distance constraint. Suppose a scenario that the vehicle traces an assigned path shown in Fig. 1. Denote s as the arc length measured along the assigned path, and q s\u00f0 \u00de is the path curvature as a function of s. Meanwhile, sf denotes the total length of the path, and tf is the total travel time of vehicle in the assigned path. After that, the time-optimal problem could be carried out by searching a velocity profile v t\u00f0 \u00de according to the following optimization problem min tf \u00f012\u00de subject to s tf \u00bc sf with s t\u00f0 \u00de :\u00bc Z t 0 s n\u00f0 \u00dedn \u00f013\u00de v 0\u00f0 \u00de \u00bc v s; v tf \u00bc v f \u00f014\u00de 0 6 v t\u00f0 \u00de 6 vmax; t 2 0; tf \u00f015\u00de aLmin 6 _v t\u00f0 \u00de 6 aLmax; t 2 0; tf \u00f016\u00de v2 t\u00f0 \u00de q s t\u00f0 \u00de\u00f0 \u00dej j 6 aNmax; t 2 0; tf \u00f017\u00de The constraint (14) is used to define the initial and final velocities during the moving of vehicle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure34-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure34-1.png", "caption": "Figure 34 . Maximum total Deformed thermomechanical coupling", "texts": [ " It is observed that the contact pressure distribution of the the pads increase in a notable way, when the thermal and mechanical aspect are coupled. In Fig 32, it is found that the presence of the slot affects negatively the strain on the disc surface, unlike the use of the double-piston. Figure 33 shows that the shear stresses and sliding distance of the inner plate are symmetrical with respect to the groove and is highest at the edges. During a braking maneuver, the maximum temperature reached on the tracks depends on the storage capacity of the thermal energy in the disc. It is seen in Figure 34, that the maximum displacement is localized on the slopes of friction, the fins and the outer ring. This phenomenon is explained by the fact that the deformation of the disc due to heat (the umbrella effect) which can lead to damage by cracking. In this case, the thermo- coupling is quite important. Thermal gradients and expansions generate thermal stresses in addition to mechanical stress. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002274_j.mechmachtheory.2019.103748-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002274_j.mechmachtheory.2019.103748-Figure6-1.png", "caption": "Fig. 6. RRRP or RRRC linkage.", "texts": [ " The revolute joint can be split up in a similar way as prismatic and cylindrical joints (see Section 3.2 ). The advantage is to avoid the effects of unknown location of the forces. The universal or Hooke joint transmits tree independent forces and one independent torque [22] . When it is moved, in the presence of friction, two addition friction torques are created. Because those torques are orthogonal, they can be modelled in the same way discussed in Section 3.3 . For the efficiency analysis, the slider-crank illustrated in Fig. 6 can be considered as being composed of four rigid links connected by three revolute joins and one cylindrical or prismatic joint. Herein, this latter pair is considered prismatic since the rotation of the pair D (see Fig. 6 ) is completely prevented by the remaining members of the kinematic chain under the rigid link assumption. Assuming the coupling D as a prismatic pair decreases the dimension d and it is more of a convenience rather than a necessity. Perfect linkage geometry is assumed implying that the actions related to any overconstraint, namely the forces parallel to z -axis and torques parallel to the plane z = 0 , are neglected. In order to keep the example short, only friction in coupling D is modelled. Normally, the location of a prismatic pair like coupling D is arbitrary", " Thus, it is convenient to locate coupling D coincident with C so that the torques at the origin produced by the vertical force in couplings C and D compensate each other. Moreover, coupling D behaves like the slider in Fig. 1 and only the first mode of operations discussed in Section 3 can take place. The position of couplings B, C and D can be expressed in terms of \u03b8 , the crank angle, and the link lengths a 1 and a 2 as: B x = a 1 cos \u03b8 B y = a 1 sin \u03b8 C x = D x = B x + \u221a a 2 2 \u2212 a 1 2 + B x 2 = a 1 cos \u03b8 + \u221a a 2 2 \u2212 a 1 2 + a 1 2 cos 2 \u03b8. (6) The motion screws of the coupling network shown in Fig. 6 belong to the 5th special 3-system of screws according to Hunt\u2019s [18] classification. By orientating the z -axis parallel to the ISA of couplings labelled A, B and C, and the x -axis parallel to the ISA of coupling D, the motions can be spanned by the three motion screws { t ; u , v }. The motion screws associated with the couplings are detailed in Table 2 . The action screws in the coupling network shown in Fig. 6 , which can be spanned by { T ; U , V }, belong to the 4th special 3-system of screws, i. e., the screws of planar statics [18] . These action screws are described in Table 3 . Note that, in Table 3 , T A is the magnitude of an external torque and U D is that of an external force, both applied to the network. U D D is the magnitude of the force due to friction given by U D D = \u2212\u03bcD sign ( V D ) V D sign ( u D ) (7) where V D is the magnitude of the vertical force transmitted by coupling D (the normal force), \u03bcD is a dimensionless coefficient of friction in the classical sense, and u D is the magnitude of translational velocity of coupling D (the speed of the slider with respect to the fixed frame). The minus sign in Eq. (7) is due to the fact that both screws $ a U D D and $ a u D are orientated from left to right. Eq. (7) is in fact the Coulomb friction model, but more complex friction model can be adopted. The motion analysis of the single loop kinematic chain machine illustrated in Fig. 6 is a trivial matter since the network unit motion matrix \u02c6 M N is equal to the unit motion matrix of the direct couplings \u02c6 M D obtained from Table 2 as \u02c6 M N = \u02c6 M D = [ t A t B t C u D t 1 1 1 0 u 0 B y 0 1 v 0 \u2212B x \u2212C x 0 ] . (8) where B x , B y , and C x are given by Eq. (6) . The null space, also called the kernel, of \u02c6 M N and \u02c6 M D is Null ( \u02c6 M N ) = \u23a1 \u23a2 \u23a3 t A C x \u2212 B x t B \u2212C x t C B x u D B y C x \u23a4 \u23a5 \u23a6 1 C x \u2212B x . (9) The solution of Eq. (1) is proportional to the sole vector of Null ( \u02c6 M N ) . Any variable listed in Table 2 can be chosen as the primary variable. Arbitrarily chosen t A , the motion screw matrix that contains one screw per column, is given by M = \u02c6 M D diag ( Null ( \u02c6 M N )) t A (10) which, using Eqs. (8) and (9) , leads to M = [ $ m A $ m B $ m C $ m D t C x \u2212 B x \u2212C x B x 0 u 0 \u2212B y C x 0 B y C x v 0 B x C x \u2212B x C x 0 ] t A C x \u2212B x . (11) The topology of the coupling network schematically shown in Fig. 6 is represented by the coupling graph of Fig. 7 . Using edges A, B, and C as branches, the fundamental cutset matrix of G C is obtained by inspection as Q = [ A B C D A 1 0 0 \u22121 B 0 1 0 \u22121 C 0 0 1 \u22121 ] . (12) In this case, any branch choice results in the same cutset matrix. For fundamental cutset matrix obtention, refer to any of the references [10,11,15,23\u201325] . The fundamental cutset matrix of action graph G A is obtained by means of column replication as Q A = \u23a1 \u23a3 A T A A U A A V A B U B B V B C U C C V C D T D D U D D V D D U D D A 1 1 1 0 0 0 0 \u22121 \u22121 \u22121 \u22121 B 0 0 0 1 1 0 0 \u22121 \u22121 \u22121 \u22121 C 0 0 0 0 0 1 1 \u22121 \u22121 \u22121 \u22121 \u23a4 \u23a6 (13) where every column of Q was replicated a number of times equal to the number of independent actions transmitted by the respective coupling (see Table 3 )" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001310_1.4029709-Figure18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001310_1.4029709-Figure18-1.png", "caption": "Fig. 18 Orbit of the rotor at bearing position and at shaft center ( 5 10 krpm\u00de", "texts": [ " The unbalance is modeled using two masses disposed at the shaft ends (node number 1 and 13) of 40gmm each. The two unbalances have the same phase shift. The rotating speed is 10 krpm and the rotor is stable in the sense of 092502-8 / Vol. 137, SEPTEMBER 2015 Transactions of the ASME Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/28/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use small perturbations [7]. The integration of Eq. (20a) was performed over 512 periods [16\u201320]. Figure 18 shows the orbit of the rotor at bearing position and at the shaft center during the 256 last periods. As shown, the orbits are complex and contain a number of closed loops. The power spectrum of the displacement in the X direction is presented in Fig. 19. In this case, the rotor describes a quasi-periodic motion as shown by the Poincar e diagrams depicted in Figs. 20 and 21. In this paper, the method proposed in Ref. [6] was enhanced by using the same set of stable poles that ensure naturally the continuity of the fluid forces when dealing with large displacements" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001935_s11668-016-0155-5-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001935_s11668-016-0155-5-Figure3-1.png", "caption": "Fig. 3 The bearing seeded faults", "texts": [ " This section covers the experimental setup used to validate the MA-based AE bearing fault diagnostic technique. Figure 2 shows the bearing test rig used to collect the AE data and conduct the bearing seeded fault test. A wide band (WD)-type AE sensor was axially mounted on the face of the bearing housing using instant glue. Type 6205-2RS steel FAG ball bearings were used for testing. Four fault types were simulated on steel bearings: inner and outer race faults, rolling element fault, and cage fault (see Fig. 3). The inner and outer race faults were generated by scratching the steel race surfaces with a diamond tip grinding wheel bit to cover the ball contact surface. The ball fault damage was created by using the grinding wheel bit to create a small dent in one of the steel balls. For the cage fault, the steel cage was cut in between two ball locations. For all seeded fault tests, the bearing seal and grease was removed and replaced following the creation of the fault. Figure 4 shows the AE data acquisition system consisting of a demodulation board, power supply, along with the function generator, and sampling device" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002701_aero47225.2020.9172614-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002701_aero47225.2020.9172614-Figure12-1.png", "caption": "Figure 12: Dimensions of a regular tetrahedron [5].", "texts": [ " The electronics used are commonly used in conventional quadrotors. The Tetracopter uses four HQProp Ethix S5 5x4x3 rotors and Emax Eco 2306 2400kv motors. The CC3D flight controller along with four HAKRC BLHeli-32 Bit 35A 2-5s electronic speed controllers are used to control the motors. To fully characterize the proposed aircraft a number of parameters must be defined starting with the shape of the frame. As stated before, the frame of the single-propeller module is in the shape of a tetrahedron as seen in Fig. 12. Using the variables in Fig. 12, the length of a side of the single-propeller submodule a is equal to 244.55mm. Given a the remaining variables seen in Fig. 12, can be defined using equations (28) to (32) [5]. x = 1 3 \u221a 3a (28) d = 1 6 \u221a 3a (29) h = 1 3 \u221a 6a (30) R = 1 4 \u221a 6a (31) \u03c6 = tan\u22121( r x ) (32) The distance from the axis of the center rotor to the outer rotors, labeled as \u03bc in Fig. 10a, is approximately 130.4mm while the distance between the planes labeled in Fig. 10b as h is 168.8mm. The total mass of the vehicle is approximately 740 g. Given the degradation in thrust due to the frame of the Tetrahedral rotor-craft, at 75% motor throttle the estimated thrust to weight ratio is 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002246_j.triboint.2020.106222-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002246_j.triboint.2020.106222-Figure2-1.png", "caption": "Fig. 2. General architecture of the whole model.", "texts": [ " In this investigation, compared with conventionally rigid pivot bearing, a kind of TPJB is introduced with disc spring set pivot design and pivot stiffness is controlled to vary from a completely soft one to an almost rigid one. In addition, the journal unbalance eccentricity ratio is varied. Journal orbit, fluid film maximum pressure, minimum thickness and maximum tilt angle are predicted and researched. A kind of TPJB with disc spring set pivot design is employed in Fig. 1 and a separate pad model with disc spring set pivot is given. The proposed model architecture [40-43] for the TPJB with disc spring pivot design consists of three following parts, as presented in Fig. 2: (1) Pad model, including a fluid film sub-model and a pad dynamic sub-model for each pad: (a) The fluid film sub-model represents the specific fluid film interposed between the journal and each pad, according to the Reynolds equation under the nonlinear regime. (b) The single pad dynamic sub-model represents the pad radial motion and rotation around its pivot. The radial motion for each pad is due to the compliance of corresponding disc spring sets. (2) Pivot model, i.e., the stiffness of disc spring pivot design, coupling the disc spring sets with the pad model" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001268_s0263574714002653-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001268_s0263574714002653-Figure6-1.png", "caption": "Fig. 6. (Colour online) Model of the spherical joint.", "texts": [ "\u03c4 ext 0 ] , (27) where f ij corresponds to the force transmitted by the j-th spherical pair of the i-th PKL. Let uij, be the unit vector of the force f ij, which can be written as uij = f ij f ij = Hij.\u03c4 ext Hij.\u03c4 ext Hij, (28) where Hij is the matrix that relates the reaction forces transmitted by the j-th spherical joint of the i-th PKL to the external load \u03c4 ext applied on the platform (Appendix B and C). http://journals.cambridge.org Downloaded: 12 Mar 2015 IP address: 131.183.72.12 Thus, the change of the pose caused by the clearance in the spherical joint is given by the following vector (Fig. 6): \u03b4sij = C1C2 = \u03b5.uij, (29) where \u03b5 corresponds to the magnitude of the vector \u03b4sij . According to Eqs. (28 and 29), the change of the pose caused by the clearance in the spherical joint can be written as \u03b4sij = Hij.\u03c4 ext\u2225\u2225Hij.\u03c4 ext \u2225\u2225 .\u03b5. (30) The local pose error caused by the clearance in the revolute pairs is determined using the model given by18. First, the reaction forces and moments \u03c4 rij transmitted by the revolute joints is equivalent to three contact forces C1,ij, C2,ij and C3,ij http://journals" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000486_s11771-019-4180-x-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000486_s11771-019-4180-x-Figure7-1.png", "caption": "Figure 7 Non-modified gear set frictionless bending stress distribution", "texts": [ " (2019) 26: 2368\u22122378 2373 Similarly, parabolic tip relief was also applied using the same normalised amount and length of modification values used for linear tip relief (see Figure 5). The maximum frictionless contact stress of the non-modified gear set was recorded on the gear tooth subsurface with a value of 1227.5 MPa. Figure 6 shows the contact stress distribution which is higher close to the contact area and decreases away from it. Similarly, the maximum bending stress was also recorded on the gear tooth root. The stress can be seen to decrease with distance away from the gear tooth fillet area as shown in Figure 7. The non-modified gear set frictionless contact stress of 1227.5 MPa was compared with the calculated values reported by PATIL et al [15] for validation purposes (see Table 4). The Hertzian contact stress was calculated as 1284 MPa and AGMA pitting resistance was calculated as 1621 MPa, while Ref. [15] reported a maximum contact stress of 1240 MPa. It can be seen that the current result agrees well with the result reported by PATIL et al [15], with a deviation of only 1.0%. Likewise, the Lewis bending stress and AGMA bending strength were calculated as 234 and 289 MPa, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001246_j.ijheatmasstransfer.2014.07.091-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001246_j.ijheatmasstransfer.2014.07.091-Figure2-1.png", "caption": "Fig. 2. (a) The physical model and coordinate system, and (b) pore formation from enclosure of the keyhole.", "texts": [ " This simple, general and flexible method has been extensively and efficiently used to investigate the complicated annular two-phase flows and their transitions from annular to slug flows. Slug flow is a liquid\u2013gas two-phase flow in which the gas phase exists as large bubbles separated by liquid slugs. This work provides a systematical and fundamental step to understand the effects of the compressible mixture gas in the keyhole on the formation and final size of a pore taking place at the base. In this study, the co-ordinate system and geometry of the keyhole are illustrated in Fig. 2(a). For convenience, the solid can be considered to move upward at a constant relative speed U relative to the liquid\u2013solid interface. In view of the incident flux, a molten pool in a thin layer is formed around the keyhole. Mixture flux is ejected at the keyhole base and entrained through the side wall. The core region thus contains a mixture comprised of vapor and droplets [20\u201324]. The keyhole is often susceptible to collapse and block the incident flux, leading to enclosure of the keyhole and formation of the pore at the base, as illustrated in Fig. 2(b). The pore at the keyhole base is approximately a sphere, characterized by an effective radius rp. Without loss of generality, the major assumptions made are the following: 1. The model is one-dimensional. This is valid for a deep and narrow keyhole produced by a high power density beam or small welding speed. 2. The model is axisymmetric in keyhole welding with a low scanning speed. The reason for this is that the time scale for keyhole collapse is around 10 3 to 10 4 s [12,13], which is much less than time scale for welding around 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001279_s0263574714000149-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001279_s0263574714000149-Figure10-1.png", "caption": "Fig. 10. Mobility calculations for overconstrained mechanisms.", "texts": [ " Moreover, for the kinematic analysis model of the four bar mechanism the mobility calculation suggest that a number of \u2211 fi = 4 constraint equations will be formulated and a DAE system with 7 equations for the dynamic model. This technique does not fail for any mechanism and demonstrates that the statement of Freudenstein and Alizade in ref. [10] according to which the mobility formula given by Eq. (6) \u201capplies to mechanisms without exception\u201d is true. Other examples of particular spatiality mechanisms are illustrated in Fig. 10. For the case of complex mechanisms (with more than two closed loops) the same KCB mobility formula and method for determining the spatiality of loops remains valid and gives correct results. Nevertheless, for multi-loop mechanisms spatiality identification, the closed loops should be open successively and not simultaneously. To justify this statement, we recall that the aim of S identification is to avoid introducing redundant constraints in the mobility calculation. In the case of multiple loops, it is possible that some motions of a certain loop are already restricted by other closed loops and therefore they cannot be counted again" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001152_s12206-014-0814-y-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001152_s12206-014-0814-y-Figure2-1.png", "caption": "Fig. 2. The acceleration testing positions and the rubbing positions.", "texts": [ " Finally the experimental rig of aero-engine forms the rotor-support-blade disk -casing system in structure. The experiment rig is shown in Fig. 1. Four rubbing screws are designed on turbine casing to carry out the rubbing experiments at the four positions. Four accelerometers are placed on the turbine casing to collect the casing acceleration signals. The rubbing positions and the installation positions of accelerometers are shown in Figs. 1 and 2. Facing to the turbine casing, the directions of rubbing positions and installation positions of accelerometers are shown in Fig. 2, which also shows the testing channels of the four accelerometers. Six independent experiments were carried out in two days, which include three experiments on 2012-5-12, and three experiments on 2012-5-1. Four different rubbing positions were seeded, which include the upper, the lower, the left and the right. The 100 fault samples including four different rubbing positions were obtained in each experiment, and each sample includes 8192 collecting points. The sampling frequency is 10 kHz. All the experiments are carried out at 1500 rpm" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.14-1.png", "caption": "Fig. 3.14 Gimbal lock: the orientation of the innermost axis cannot be reoriented in the direction of the dotted arrows", "texts": [ " Angles from rotation sequences that involve all three axes (x-y-z, y-z-x, z-x-y, x-z-y, z-y-x, y-x-z) can be either called \u201cTait\u2013Bryan angles\u201d, in honor of the Scottish mathematical physicist Peter Tait (1832\u20131901), who was\u2014 together with Hamilton\u2014the leading exponent of quaternions, and the Welshman George Bryan (1864\u20131928), the originator of the equations of airplane motion or \u201cCardan angles\u201d, after the Renaissance mathematician, physician, astrologer, and gambler Jerome Cardan (1501\u20131576), who first described the cardan joint which can transmit rotary motion. And angles that have the same axis for the first and the last rotation (like the Euler sequence above) are called \u201cproper Euler angles\u201d (z-x-z, x-y-x, y-z-y, z-y-z, x-z-x, y-x-y). Consider tracking a helicopter flying from the horizon toward an aerial gun, as indicated in Fig. 3.16. The helicopter flies toward the gun site and is tracked by the gun in elevation ( ) and azimuth ( ). When the helicopter is immediately above the gun site, the aerial gun is in the orientation indicated in Fig. 3.14. If the helicopter now changes direction and flies at 90\u25e6 to its previous course, the gun cannot track this maneuver without a discontinuous jump in one or both of the gimbal orientations. There is no continuous motion that allows it to follow the target\u2014it is in \u201cgimbal lock\u201d. Note that even if the helicopter does not pass through the gimbal\u2019s zenith, but only near it, so that gimbal lock does not occur, the system must still move exceptionally rapidly to track the helicopter if it changed direction, as it rapidly passes from one bearing to the other" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000353_iemdc.2019.8785381-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000353_iemdc.2019.8785381-Figure2-1.png", "caption": "Fig. 2. Final field solution of the iterative procedure for the proper rotor current computation, according to the inverse-\u0393 model of IM.", "texts": [ " An important tool, associated to this technique is the equivalent rotor three-phase winding, that allow the cage to be treated as a common winding. In this method, the stator current vector is considered to be fixed in the chosen dq reference frame. According to (5), the rotor current imposes its magneto-motive force along the qaxis, and an iterative loop is meant to get the proper amplitude of the current in order to have \u03bbrq = 0. In original method, rotor current is supposed to be sinusoidally distributed along the rotor periphery. The flux map, after the orientation, is shown in Fig. 2. It can be noticed that the flux lines within the rotor are almost parallel to d-axis, which means that the component \u03bbrq is almost zero. Once the RFO field solution is achieved, the motor performance can be derived, using the model equations of torque and slip: Tdq = 3 2 p (\u03bbsdisq \u2212 \u03bbsqisd) ; s = Rr \u03c9sLr isq isd (6) where \u03bbrd and \u03bbrq are directly derived from the RFO field solution, \u03c9s is the electric angular frequency and p is the number of pole pairs. C. Inclusion of Skewing The rotor bars skewing is an effective solution to reduce the torque ripple and additional under-load Joule losses in the cage and for avoiding synchronous torques" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001977_j.ijfatigue.2016.09.006-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001977_j.ijfatigue.2016.09.006-Figure10-1.png", "caption": "Fig. 10. Discrete models used for Kt determination : (a) for test pieces with V-notch; (b) for test pieces with U-notch.", "texts": [ "015 0.005 0.025 0.036 0.012 Fig. 4. Parent material test pieces for tests under tension-compression, (mm). Fig. 7. Welded joints for tests under tension-compression and oscillatory bending, where t = 10 mm (tension-compression) and t = 30 mm (oscillatory bending), (mm). oscillatory bending Kt-w = 2.06 (according to [25]). In order to correctly choose geometric parameters of U and V notches, analysis of stress fields was conducted by means of the Finite Element Method for linear-elastic body model. Fig. 10 presents discrete models used in analysis. The undertaken finite element analysis allowed to investigate the notch geometry in order to find geometric parameters that Please cite this article in press as: \u0141agoda T et al. Investigation on the effect of joints. Int J Fatigue (2016), http://dx.doi.org/10.1016/j.ijfatigue.2016.09.006 result in Kt-w (U, V) reflecting Kt-w of welded joints. In this way, appropriate values of notch root radius are adopted in the test pieces shown in Figs. 8 (bending) and 9 (tension-compression)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure20-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure20-1.png", "caption": "Fig. 20. Eye-connection plates.", "texts": [ " It should be noted that the test rig design procedures discussed here are for testing nodes under individual loading case. Similar procedures can be applied to the design for combined loading case. The first step is to calculate the eccentricities and apply the axial loads at the eccentric centres of the connecting faces (Fig. 19). In this example, constant eccentricities of 30 mm are applied in both in-plane and out-of-plane directions. In the second step, eye-connection plates are designed to be vertically mounted on the connecting members as shown in Fig. 20. As discussed before, the positions of bolt holes on these plates should be determined with the consideration of the bolt looseness, so that after initial movement, the bolts should be placed in the right positions as they are designed for. As can be seen in Fig. 20, the vertical plates are parallel to the connected members\u2019 centre-lines. In the third step, the rigid rotating plates are designed to transform the applied vertical load (Vi) to the design load (Fi) as shown in Fig. 21. The last step is to design a set of balanced beams for combining vertical loads to obtain one resultant load (F) as shown in Fig. 22. In this study, an innovative test rig is designed to test two different types of nodes. By slightly adjusting the setup, the rig can be used to test structural nodes under various loading conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001692_1.4943902-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001692_1.4943902-Figure2-1.png", "caption": "FIG. 2. Typical shape of a turbine blade with indication of sections with heat accumulation (highlighted red).", "texts": [ " In order to suppress the formation of hot and cold cracking during the build-up phase of the components, the implementation of tailored temperature cycles by means of additional heat sources is unavoidable in most cases. To ensure such customized temperature regimes, the typically used inductive heating systems need to be extended by additional process monitoring equipment. This becomes even more obvious when considering a typical shape of a turbine blade and imaging the change in heat distribution during the movement of the laser (see Fig. 2). Due to heat accumulation causing an increased viscosity in the melt pool, the layerwise build-up will be inhibited. Hence, an absence of heat might cause an underrun of the minimum temperature. So the fundamental challenge is the precise balancing of the induced heat by inductive heating and the laser beam, as well as considering the heat distribution as a function of the geometry and the surrounding atmosphere. The Fraunhofer Institute for Material and Beam Technology Dresden (IWS) has developed such a temperature monitoring and control system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001220_j.triboint.2014.10.021-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001220_j.triboint.2014.10.021-Figure1-1.png", "caption": "Fig. 1. Schematic of Toyota 1NZK engine end pivoted finger follower valve train.", "texts": [ " Engine valve train is one of the most complicated tribological mechanisms in the engine which often operate in severe operating conditions due to continuous variations in the contact loading, higher operating temperatures and varying nature of cam geometrical profile affecting the lubricant entrainment velocity and thus the lubrication mode. In addition, engine valve train has witnessed numerous design changes over the years to improve performance in the form of higher spring stiffness, multi valves system, variable valve and cam timing. These design factors along with the use of low viscosity lubricant have pushed the valve train mechanism to operate in much harsh operating conditions than before. In automotive sector, there is a wide spread use of end pivoted roller follower valve train (Fig. 1) in the modern passenger automobiles due to their improved power output, better fuel economy, compact cylinder head and smoother valve operation. Many researchers [1\u20133] have reported substantial reduction in the power loss by using the roller followers instead of sliding tappets. The improved lubrication condition at cam/roller contact due to roller rotation was recognized by Jonmin and Taylor [4]. In roller follower valve train, roller rotation plays an important role to reduce the friction, wear and minimizes the chances of fatigue failure by even distribution of wear" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002488_tmag.2020.2997759-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002488_tmag.2020.2997759-Figure2-1.png", "caption": "Fig. 2: Schematic of Optimized Levitator", "texts": [ " The levitator setup consists of two cylindrical coils carrying time-varying sinusoidal current. The coil closer to the centre (i.e. the primary coil) is used for the generation of the primary vertical force. The coil closer to the outer edge (i.e. the secondary coil) is utilized for generation of stabilizing lateral forces. These coils are embedded within a ferrite core block. The various dimensions of the system are obtained through isolated optimization of the parameters. The schematic for the levitator setup are shown in the Fig. 2. The technique presented has several significant advantages over conventional techniques. The levitated part serves as an initial base for building the part. Ideally, the base would end up being a part of the final component. Fig. 2 shows the use of multiple material jets to build the part. This idea is further supported by the research presented in [9]. Thus, the system allows for multi-directional deposition, thereby improving the efficiency of building the part. In addition to the use of multiple material jets, the levitator setup is also capable of producing the torque necessary to flip Authorized licensed use limited to: University of Wollongong. Downloaded on May 30,2020 at 01:10:09 UTC from IEEE Xplore. Restrictions apply" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000701_humanoids43949.2019.9035038-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000701_humanoids43949.2019.9035038-Figure4-1.png", "caption": "Fig. 4: Estimation of the wrench at the end effector based on the wrench measured with the F/T sensor. It is worth to note that inertial forces are neglected. The end effector is assumed to be steady or moving slowly if it is pushing.", "texts": [ " Let us use a Low-Pass filter on the measured signals and name the resulting signals as sFs = [ sfTs snTs ]T . Then, the equivalent wrench with respect to a of the end effector, expressed in the world frame, can be estimated as: FEF/a = [ fEF nEF/a ] = [ Rs ( sfs)\u2212mEg Rs ( sns)\u2212 S ( rg/a ) mEg ] . (15) Here, mE is the equivalent mass of the portion of the end effector \u201cafter\u201d the F/T sensor. g is the vector of gravity (pointing downwards). rg/a is the position of the center of gravity of this portion with respect to a. See Fig. 4. Finally, notice the minus sign multiplying the positive PD gains of (11). The directions of the bases \u03b2c,i,1 are chosen such that they generate forces exerted by the environment on the link. As such, they are directed towards the link. This is also the case of the wrench measured by the F/T sensor. On the other hand, we want to produce accelerations based on the forces that the link should be exerting on the environment. These ones should drive the link against the environment. The minus sign is used to invert the direction of the produced acceleration" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000853_978-3-642-40066-7_10-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000853_978-3-642-40066-7_10-Figure5-1.png", "caption": "Fig. 5 Illustration of Lemma 1 for spanner property of Yao structure", "texts": [ " Yao graph [107] is originally proposed for construction of high-dimensional MST, but has been recently used for topology control in 2D ad hoc and sensor networks [58, 59]. In 2D, Yao graph Y Gk is defined as follows. At each node u, any k equally-separated rays originating at u define k cones. In each cone, choose the shortest edge uv among all edges emanated from u, if there is any, and add a directed link \u2212\u2192uv. In [58, 59], Li et al. proved that 2D Yao graph is an energy spanner of UDG. Their proof of the spanner property is based on the induction of link length. As shown in Fig. 5, for each removed link uv /\u2208 Y Gk , they proved the energy consumed by a shorter link uw and a path from w to v is within constant times of the energy consumed by link uv. The key result of their proof can be summarized by the following lemma: Lemma 1. [58, 59] The energy stretch factor of the Yao-based graph is at most 1 1 \u2212 (2 sin \u03b4 2 )\u03b2 , if for every link uv that is not in the final graph, there exists a shorter link uw in the graph and \u2220vuw < \u03b4, where \u03b4 is a constant smaller than \u03c0/3, as shown in Fig.5. Clearly, if k > 6, 2D Yao graph has its energy stretch factor bounded by 1 1 \u2212 (2 sin \u03c0 k )\u03b2 . 3D Yao structures can use certain types of 3D cones to partition the transmission region of a node (which is a sphere), and inside each 3D cone the node only keeps a link to the nearest neighbor. If the number of such 3D cones is bounded by a constant k, 3D Yao structures can bound the node out-degree by k. However, it is hard to define the partition boundary of Yao structure of a node in 3D. Notice that a disk in 2D can be easily divided into k equal 2D cones which do not intersect with each other, but in 3D case, it is hard to divide a sphere into k equal 3D cones without intersections among each other" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002640_j.isatra.2020.07.039-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002640_j.isatra.2020.07.039-Figure1-1.png", "caption": "Fig. 1. The CWRU test rig.", "texts": [ " In Section 5, kernel functions and other details of the IIF are analyzed. Eventually, Section 6 gives the conclusion. 2. Construction procedure of training dataset A brief introduction of the public datasets is presented in this section. Then, on the basis of analyzing the fault impulse signal, a fault-impulse sample dataset is constructed to prepare for the subsequent work. 2.1. Data source: the CWRU dataset The investigation is carried out with the Case Western Reserve University (CWRU) bearing datasets [20]. The test rig is demonstrated in Fig. 1, which consists of a AC motor, a torque sensor and a dynamometer. The bearing type of the motor is 6205-2RS JEM SKF. Table 1 shows the relevant bearing specifications. Singlepoint defect of bearing was machined, which was reinstalled into the testing platform. More details of the CWRU dataset can be inquired from data center website [20]. 2.2. Fault-impulse sample dataset setup The CWRU dataset is suitable for constructing the fault impulse dataset, because it has relatively obvious fault impulse waveforms, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000094_acs.2979-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000094_acs.2979-Figure1-1.png", "caption": "FIGURE 1 Reference frames of the medium-scale unmanned autonomous helicopter", "texts": [ " According to the flight dynamics and aerodynamics,34,35 this section displays the 6-DOF nonlinear dynamics model of medium-scale UAH with rotor flapping dynamics, actuator faults, and external unknown disturbances. The detailed processes are elaborated as follows. In order to indicate the movement of 6-DOF medium-scale UAH more intuitively, two reference frames are introduced firstly. Let \u211ce = {Oe, X\u20d7e, Y\u20d7e, Z\u20d7e} denote the inertial frame and it is fixed to a certain point of the earth. Let \u211cb = {Ob, X\u20d7b, Y\u20d7b, Z\u20d7b} be the body axis frame and it is fastened to the Center of Gravity (CG) of the UAH. The corresponding direction of each vector is revealed in Figure 1.26 Taking the longitudinal and lateral flapping dynamics of the main rotor into account, the 6-DOF nonlinear rigid body dynamic model of medium-scale UAH with external unknown disturbances can be expressed by the Newton-Euler equations as follows36,37: P\u0307 = V mV\u0307 = RF + mge + d1(t) \u039b\u0307 = H\u03a9 J\u03a9\u0307 = \u2212\u03a9 \u00d7 J\u03a9 + \u03a3 + d2(t) \ud835\udf0fea\u0307 = \u2212a \u2212 \ud835\udf0feq + AlonTa \ud835\udf0feb\u0307 = \u2212b \u2212 \ud835\udf0fep + BlatTb, (1) where m is the gross mass, g is the gravitational acceleration, and e = [0, 0, 1]T. J = diag{Jx x, Jy y, Jz z} is the inertia matrix, and P = [Xg,Yg,Zg]T and V = [u, v,w]T represent the position vector and velocity vector in the inertial frame, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000436_acc.2019.8815013-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000436_acc.2019.8815013-Figure2-1.png", "caption": "Fig. 2. Tilt-Rotor Quadcopter UAV", "texts": [ " Wind disturbances as step inputs have been studied on UAVs by some researchers [7] [15]. However, wind disturbances have not been studied on tilt-rotor quadcopters and how the servo motor functionality of these quadcopters can help the vehicle to achieve its desired state. The focus 978-1-5386-7926-5/$31.00 \u00a92019 AACC 2005 of this paper is to explore the dynamics of the previously studied tilt-rotor quadcopters (based on hardware designed and fabricated in our laboratory) with capabilities of rotating their motor-propellers as shown in Figure 2 and provide insights on the stability characteristics while hovering at a certain desired attitude/orientation. Also, a novel sliding mode controller has been designed that uses the dynamics of this fully actuated system to obtain control inputs that enable the UAV to achieve a desired orientation and position at a given point of time with minimal errors. The robustness of this controller is then demonstrated under various wind disturbance conditions which were obtained by performing CFD simulations on the tilt-rotor quadcopter CAD model. In the next section, the dynamics of the tilt-rotor quadcopter are presented followed by which a non-linear sliding mode controller is derived for this UAV. Numerical simulation results for various scenarios are then presented followed by conclusions. The dynamics of the TRQ system is studied in this section. The free body diagram, highlighting the forces and moments acting on a tilt-rotor quadcopter, is shown in Figure 2. The world-frame (W ), denotes the inertial frame and the motion of tilt-rotor quadcopter is considered with respect to it. Similarly, the body-frame (B), is a frame attached to the center of mass of the vehicle. In Figure 2, the XB and YB axes are in line with the quadcopter\u2019s arm with XB pointing forward and YB pointing left. The vehicle is a \u2018+\u2019 configuration quadcopter with positive XB defined as forward direction as shown in the figure. Thrust force, generated by each motor-propeller set and the moment-reaction produced by a motor-propeller set on the vehicle (directed opposite to the direction of rotation of the propeller) is also shown in Figure 2. There is a force from each propeller acting in direction perpendicular to the rotor plane. It can be seen that the motor-propeller sets have tilt functionality about their respective quadcopter arm. The planes with dashed lines are the original rotor planes with tilt angles \u03b8i = 0. Similarly, the planes with the rigid lines are the tilted rotor planes for the corresponding propellers. The translational equations of motion along the X,Y and Z directions in the world frame are: m x\u0308y\u0308 z\u0308 = Rw b F2s\u03b82 + F4s\u03b84 \u2212F1s\u03b81 \u2212 F3s\u03b83 F1c\u03b81 + F2c\u03b82 + F3c\u03b83 + F4c\u03b84 + m 0 0 \u2212g (1) where m is the total mass of the quadcopter, g is the acceleration due to gravity, \u03b8i, (i = 1, 2, 3, 4) are the servo motor tilt angles, c\u03c8 and s\u03c8 denote cos(\u03c8) and sin(\u03c8) respectively, and, Fi = kf\u03c9 2 i (i = 1, 2, 3, 4) is defined as the forces generated by the four rotors, where \u03c9i is the angular velocity of the ith rotor and kf is the thrust coefficient" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002506_012078-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002506_012078-Figure2-1.png", "caption": "Figure 2. The relationship between the lateral force and the slip angle for various values of the curvature of the wheel treadmill and the traction coefficient (longitudinal force)", "texts": [], "surrounding_texts": [ "Most modern wheeled-vehicle tires have a treadmill (tread) that is nearly cylindrical in shape. The contact patch of such tires with a solid supporting surface is close to rectangular. Along with these, there are tires in which (in the inflated state) the tread has a convex shape (toroidal). As shown in [1], the toroidality (convexity) of the treadmill of the wheel affects the process of its linear rolling only with small longitudinal tangential forces acting in the contact area. At the same time, the question of the influence of the toroidalness of the treadmill of an elastic wheel on its rolling with a slip and along a curved path remains unclear." ] }, { "image_filename": "designv11_22_0003684_j.matpr.2021.05.466-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003684_j.matpr.2021.05.466-Figure9-1.png", "caption": "Fig. 9. 3-D Model of Elliptical Head Cylinder.", "texts": [], "surrounding_texts": [ "Fig. 7 depicts A circular AB thin-walled tank under internal pressure. On the tank\u2019s wall is depicted a stress element whose faces are perpendicular and parallel to the tank\u2019s axis. The normal stresses r1 and r2 acts on the element\u2019s side face. Because of the symmetry of the vessel and its loading, shear stresses do not act on the faces. As a result, stresses r1 and r2 are primary stresses. Due to their directions, stress r1 is known as the hoop stress or circumferential stress, and stress r2 is known as the axial stress or longitudinal stress [31,32]. The following equations are based on several assumptions for circumferential and longitudinal stress [33]: 1. Plane sections remain plane. 2. r/t 10 where t is uniform and constant. Fig. 6. Plain For 3. The material is linear-elastic, isotropic, and homogeneous in nature. 4. There will be no variation in stress distributions across the wall thickness. 5. The fluid\u2019s weight is considered negligible." ] }, { "image_filename": "designv11_22_0002258_s00170-020-04953-3-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002258_s00170-020-04953-3-Figure8-1.png", "caption": "Fig. 8 Gaining the points data", "texts": [ " When the actual tooth surface is within the tooth body of the theoretical tooth surface, the deviation is negative, as shown in the blue line; when the actual tooth surface is outside the tooth body of the theoretical tooth surface, the deviation is positive, as shown in the red line. The tooth surface accuracy is described by the maximum deviation value and the average deviation value, which are defined as: \u0394 f max \u00bc max i\u00bc1 imax max j\u00bc1 jmax \u0394f\u00f0 \u00de ij\u00f0 \u00de n o( ) \u00f020\u00de \u0394 f \u03a3 \u00bc \u03a3 i\u00bc1 imax \u03a3 j\u00bc1 jmax \u0394f\u00f0 \u00de ij\u00f0 \u00de n o imax jmax \u00f021\u00de The gear measurement center is used to measure the planar envelope hourglass worm tooth surface, and the measurement diagram is shown in Fig. 8. Using the scanning function of the gear measurement center to collect data, the specific steps are as follows: (1) The planar envelope hourglass worm is positioned radially through the center holes by the upper and lower centers of the gear measurement center. (2) The probe is located at the reference position\u2460, and the coordinate values can be represented as (xc (I), yc (I), zc (I)), while the distance between the reference and the throat center is used to perform axial positioning. (3) Placing the probe on the YOZ plane, that is, the X axis coordinate is 0 and the X axis is locked", " Equation (6) shows that the theoretical probe center surface is in the theoretical coordinate system \u03c31, but the measured point data are in the instrument coordinate system \u03c3c. In order to compare the measured point data and the data of theoretical Table 2 Tooth surface deviation of tooth surface Deviation value \u0394fmax (\u03bcm) \u0394f\u03a3 (\u03bcm) Tooth surface I 17.3 8.4 Tooth surface II 22.7 9.1 probe center surface, it should transformation them into the same coordinate system, and the transformation method is shown in Fig. 9. Based on the reference plane I in Fig. 8, the measured point data in instrument coordinate system \u03c3c can be transformed into the theoretical coordinate system \u03c31 by translation and rotation. The coordinate values of measured points can be represented in theoretical coordinate system \u03c31 as follows: x a\u00f0 \u00de 1 k\u00f0 \u00de \u00bc y k\u00f0 \u00de c sinc k\u00f0 \u00de c cos\u03c6c \u00fe y k\u00f0 \u00de c cosc k\u00f0 \u00de c sin\u03c6c y a\u00f0 \u00de 1 k\u00f0 \u00de \u00bc \u2212y k\u00f0 \u00de c sinc k\u00f0 \u00de c sin\u03c6c \u00fe y k\u00f0 \u00de c cosc k\u00f0 \u00de c cos\u03c6c z a\u00f0 \u00de 1 k\u00f0 \u00de \u00bc z k\u00f0 \u00de c \u2212 z I\u00f0 \u00de c \u2212h 9>>>= >>; \u00f022\u00de Here, h is the distance between the reference plane and the center of hourglass worm, and \u03c6c is the match angle", " The adjustment movement of this hourglass worm grinding machine are the A axis and the Y axis. The A axis is the adjustment motion of the grinding wheel angle, while the Y axis is the height adjustment motion of the grinding wheel. The geometry parameters of the planar enveloping hourglass worm is shown in Table 1. Based on the kinematics relation (12), the tooth surface of planar enveloping hourglass worm is ground as shown in Fig. 12. Based on the above measurement principle and method as shown in Fig. 8, the planar envelope hourglass worm sample was tested at the Klingelberg P26 gear measurement center, as shown in Fig. 13. And based on the data processing flowchart of Fig. 10, the deviation value of the planar enveloping hourglass worm tooth surface can be solved and shown in Table 2, and the normal tooth thickness at its throat is 3.517 mm. The maximum deviation value \u0394fmax and the average deviation value \u0394f\u03a3 of surface I are 17.3 \u03bcm and 8.4 \u03bcm, while surface II are 22.7 \u03bcm and 9.1 \u03bcm, and the accuracy of this planar enveloping hourglass worm tooth surface is up to level 5 in GB/T 16445-1996" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003573_j.jmrt.2021.04.034-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003573_j.jmrt.2021.04.034-Figure3-1.png", "caption": "Fig. 3 e Laser cladding square structure specimen and its biological prototype.", "texts": [ " So the distance is between 2 mm and 8 mm between two adjacent tracks. Parallel structure. The veins of the canna are approximately parallel, and this structure is resistant to cracks caused by heat. Laser cladding prepared a similar structure surface, the purpose is to retard the spread of thermal cracks, as shown in Fig. 2. Square structure. It is simplified based on the nacre of the shell. The nacre is composed of 5% organic matter and 95% aragonite in a brick wall structure. Laser cladding is used to prepare a square structure surface, as shown in Fig. 3. Triangle structure. The structure is simplified from the wings of a dragonfly, and the area ratio is less than 1:8 between the veins and the membrane. Based on the laser cladding process, a triangular bionic structure surface is prepared as shown in Fig. 4. The thermal fatigue characteristic test was carried out on the thermal fatigue testing machine. The test piece was heated to 800 C, and then it was cooled to 20 C in water. The surface cracks and morphology of the test piece were observed after 300 cycles" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000623_0954406219893721-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000623_0954406219893721-Figure2-1.png", "caption": "Figure 2. Trajectory of contact point of HCR internal gear transmission.", "texts": [ " Among them, the comparison of the involute internal gear and the novel HCR internal gear has been made in these sections. Finally, experimental verification of contact ratio of involute internal gear drive is completed in the penultimate section and the final section draws a conclusion. Tooth profile equation of the novel HCR internal gear The conjugate tooth profiles of gears can be formed by the given path of contact.37 It is assumed that the internal gear pair rotates in the counterclockwise direction as shown in Figure 2. The pinion is centered on O1, and the internal gear is centered on O2. The pitch radius of pinion and internal gear is r1 and r2, and the tip radius of pinion and internal gear is ra1 and ra2, respectively. The point of tangency of two pitch circles is P0, and the intersection point on the right side of two tip circles is P1. The perpendicular bisector of line P0P1 intersects line O1O2 at point O 0, and then take a point O00 on line O2P0 and make the line O00P0\u00bc (r2 \u2013 R). The circular arc P0P2 is formed by taking O\u2019\u2018 as center and line O00P0 as radius, and the circular arc P0P1 is formed by taking O0 as center and line O0P0 as radius, which the action arc P0P2 and P0P1 are used as the action arcs during the engaging-in process and engaging-out process" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002371_j.optlastec.2020.106206-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002371_j.optlastec.2020.106206-Figure14-1.png", "caption": "Fig. 14. 3D modeling image showing of sub specimens of tensile testing sample by using laser beam welding.", "texts": [], "surrounding_texts": [ "The tensile studies were conducted at the as weldments, dissimilar weldments and PTAC coated weldments. The mechanical properties and tensile failure photo can be seen in Fig. 12. All the specimens were tested at the room temperature. Tensile photographs of Titanium alloy Grade 2 and Ti 6Al-4V alloy, dissimilar materials and PTAC coated weldments were analyzed and concluded that all the tensile failure occurred in the heat affected zone. Additionally, the finite element analysis clearly indicated that the failures occurred at the heat affected zone and fusion zone. The butt joint strength of Titanium alloy Grade 2 and Ti 6Al-4V alloy were found to be 251 MPa and 262 MPa. The sub specimens and tensile standard specimens were obviously indicated that both of the samples reflected the same results as shown in Figs. 13 and 14. Pankaj et al. suggested that the hardness value was high in the fusion zone. Increase of laser beam and welding speed, the failures happened in the fusion zone [14]. The finite element analysis discussed with elasticity concentration of the titanium materials and tensile failures and the result showed that the failure occurred in the fusion zone and the brittle fractures were observed. The experimental results are well in accordance with the finite element model analysis. The microstructure studies showed that the coarse grain boundary directly rehabilitated in the \u03b2 phases. Due to low heat input at the fusion zone. The fast cooling rate disturbed the fusion zone and the grain boundary and coarse boundary were disintegrated and \u03b1 + \u03b2 phases were formed. However, the grain boundary growth was slightly decreased in the heat affected zone and the \u03b1 + B phase was not changed into \u03b1\u2019 martensite phases. Nevertheless the hardness value was observed as decreasing in the heat affected zone, all the failures occurred at the fusion zone because of presence of more amounts of macropits in the fusion zone. The low heat input directly affected the fusion zone, coated samples and the coarse grain boundaries were transformed into the \u03b1, martensite boundary. Owing to the application of plasma arc cladding coating in the fusion zone, the TiNi and NiCrFe boundaries were formed at the outer layer of the structure. The coated samples exhibited the elongation of 14% superior than that of the titanium materials. The NiCrBFe coating methods displayed no crack and holes at the fusion zone, owing to the transition of \u03b2 phases into NiCr phases because of low heat input. The corrosion rates were augmented in the fusion zone, heat affected zone of coated sample material and parent metals and the formation of CrN and Cr2N was not observed." ] }, { "image_filename": "designv11_22_0001352_s0025654414010026-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001352_s0025654414010026-Figure1-1.png", "caption": "Fig. 1.", "texts": [ "8) that \u0394A is always greater than zero, because the integrand is positive; i.e., the thermal energy losses in an actuator without a transmission mechanism are always greater than the thermal losses in the presence of a transmission mechanism between the actuator and the executive mechanism. The first mechanism under study consists of slider 1 of mass m to which the working resistance force F is applied and weightless roller 2 ensuring the interaction between the connecting rod and weightless driving cam 3, whose coordinate is described by the function \u03b8 (Fig. 1). The slider has mass m, and the driving cam ensures its motion according to the law x = a sin \u03b8, (6.1) where a is a constant. The slider is subjected to the working resistance force F = { F0, \u2212\u03c0/2 \u2264 \u03b8 \u2264 \u03c0/2, 0, \u03c0/2 \u2264 \u03b8 \u2264 3\u03c0/2, (6.2) whose mean value is F\u0304 = F0/2, which corresponds to aworking stroke under the load and to idle running of the unloaded mechanism. MECHANICS OF SOLIDS Vol. 49 No. 1 2014 Then the generalized force (the torque on the cam driving shaft) has the form Q(\u03b8) = { F0a cos \u03b8, \u2212\u03c0/2 < \u03b8 \u2264 \u03c0/2, 0, \u03c0/2 < \u03b8 \u2264 3\u03c0/2, Q\u0304 = 1 2\u03c0 3\u03c0/2\u222b \u2212\u03c0/2 Q(\u03b8) d\u03b8 = 1 2\u03c0 \u03c0/2\u222b \u2212\u03c0/2 F0a cos \u03b8 d\u03b8 = F0a 2\u03c0 sin \u03b8 \u2223\u2223\u03c0/2 \u2212\u03c0/2 = F0a \u03c0 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002150_mfi.2016.7849524-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002150_mfi.2016.7849524-Figure1-1.png", "caption": "Fig. 1. The Falcon V5 platform CAD model with propulsion units\u2019 configuration, CW-1,4,5,8 and CCW-2,3,6,7", "texts": [ " Conclusions and plans of further research are described in the section VI. The features of the flying platform used in experiments, applied propulsion\u2019s configuration, system of symbols describing coaxial configuration and the test bench used for data acquisition are provided in the following subsections. Falcon V5 platform was designed in Institute of Control and Information Engineering at Poznan University of Technology by PART* research group. The robot is built in X8 configuration. It consists of four coaxial propulsion drives presented in Fig. 1. Propulsion units can be divided into two types, i.e. CW-CCW and CCW-CW, with rotors rotating clockwise (CW) or counter-clockwise (CCW). 978-1-4673-9708-7/16/$31.00 \u00a92016 IEEE 418 In order to collect data for following analysis, experiments were performed on the test bench described in [15], with modified version of controller and software. Thanks to recent changes, the measurement circuit is more compact and reliable. All components of the mentioned device are showed in Fig. 2. Every propulsion unit consists of two MN3110 BLDC motors produced by RC Tiger Motors with 10\u201d carbon fiber propellers" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002865_tro.2020.3031885-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002865_tro.2020.3031885-Figure6-1.png", "caption": "Fig. 6. VT Fabrication process. First, the outline of the unfolded VT drawing, the needle hole, and the creases were precisely laser-drilled on a PVC film. Then, the carved film was folded along the creases to make the three-dimensional VT structure. After folding the film, the VT surface is silicon-coated to prevent undesired air leakage.", "texts": [ " As the four-bar linkage rotates as \u03b8f , the length of the routed tendon (blue line in the figure) changes \u0394L = {r1(\u03b81 + \u03b8f ) + r2(\u03b82 \u2212 \u03b8f ) + r3(\u03b83 + \u03b8f ) + l1 + l2 + l3} \u2212 {r1\u03b81 + r2\u03b82 + r3\u03b83 + l1 + l2 + l3} = \u03b8f (r1 \u2212 r2 + r3) (1) where \u0394L is the change in the tendon length, ri is the radius of an ith hinge, \u03b8i is the initial wind angle of an ith hinge, and li is the distance from the ith hinge to an (i+ 1)th hinge (in the case of l3, distance to the VT tip). As the four-bar linkage rotates, the length of the routed tendon can change and undesired movement of the origami actuator can be generated. This can be eliminated by choosing the radius of the hinges as follows: r2 = r1 + r3. (2) In the proposed corneal suturing robot, the radius of each hinge is set to be r2 = 4 mm and r1 = r3 = 2 mm. To fabricate the small and complex three-dimensional (3-D) VT structure, origami-based design, and pico-second laser drilling were used (see Fig. 6). First, the outline of the unfolded VT drawing, the needle hole, and the creases were precisely laser drilled on a PVC film of 250 \u03bcm thickness. During the machining, the maximum tolerance of the carved end-tip was controlled to be less than \u00b110 \u03bcm for precise tissue manipulation. Then, the carved film was folded along the creases to make the 3-D VT structure. After folding the film, the VT surface is silicon-coated to prevent undesired air leakage. As a silicon coating material, SILGUARD 184 Silicone Elastomer Kit from DOW CORNING is used" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003627_icedme52809.2021.00056-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003627_icedme52809.2021.00056-Figure10-1.png", "caption": "Figure 10. Marine wall-climbing robot based on precise control of electromagnetic adsorption", "texts": [ " In this case, electromagnetic adsorption was introduced, which was originally used in foot climbing robots. With the maturity of electromagnetic technology, it is gradually applied to crawler wall-climbing robots, breaking through the technical field of magnetic adsorption wall-climbing robots dominated by permanent magnets. In China, the Naval University of Engineering developed the first domestic marine wall-climbing robot based on precise control of electromagnetic adsorption in 2019, as shown in figure 10. Install the electromagnets on the crawlers on both sides of the body, and each adsorption unit is controlled separately. Then the master controller controls each adsorption unit to accurately control the magnetic force, making it suitable for a variety of working methods. And when crossing obstacles, it can provide a larger adsorption force than that required for a flat wall, which greatly improves the safety of its operation. However, since each adsorption unit needs to be controlled separately, it also increases the difficulty of control" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure1-1.png", "caption": "Fig. 1. Single-loop mechanism with n helical joints (nH mechanism).", "texts": [ " Two special cases are shown for which the fulfilment of the 1st- and 2nd-order mobility conditions is already sufficient for the fulfilment of all higher-order mobility conditions. Subsequently an algorithm is presented that numerically delivers a finite order of the mobility conditions that is sufficient for the finite mobility of the nH mechanism. The formulations of the mobility conditions are based on screw theory, whereby the motor notation is used that is briefly summarized in the appendix. 2. First-order kinematics of a nH mechanism For a general single-loop mechanism with n helical joints (nH mechanism) and bodies according to Fig. 1 the 1st-order closure conditions are formulated for the subsequent mobility analysis. For this purpose the loop in Fig. 1 is opened by duplicating body n into bodies 0 and n yielding an open nH chain with base body 0 and end body n. 2.1. First-order kinematics of an open nH chain The relative screw motions of the bodies of the open nH chain are described by the relative helical joint angles qi, i = 1, . . . , n around the screw axes of the helical joints that are expressed in terms of normalised screw coordinates written in the motor representation, refer to the Appendix, a\u0302i \u2261 [ ai aei ] = [ ai r\u0303iai + hiai ] , i = 1, " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000035_978-3-319-96968-8_13-Figure13.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000035_978-3-319-96968-8_13-Figure13.1-1.png", "caption": "Fig. 13.1 Schematic representation of interaction of material", "texts": [ " On the basis of the properties of the materials, selection of laser parameters should be made so that the laser can be used for production of variety of medical devices. A considerable quantity of energy in a small predetermined area is produced by laser. At the time of a beam of photons entering into a material surface, transfer of energy appears for intensification of the vibration of the molecules, causing heating to be macroscopic phenomena. Through heating, melting, or vaporizing in selected areas the focused energy achieves in a greater way than by any other process. A schematic representation of surface interaction with laser is shown in Fig. 13.1. There are significant contrasting achievements of lasers as it can perform melting, annealing, or vaporization of the material, all of which are permanently unique. Modification of the optical appearance of a surface, which a laser beam heats, is the principle of laser marking. A variety of mechanisms are there for the occurrence of this following are the two broad categories for such mechanism: (a) Marking by way of material removal from the upper surface of material and (b) Marking by material surface modification" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002865_tro.2020.3031885-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002865_tro.2020.3031885-Figure7-1.png", "caption": "Fig. 7. Illustration of the simulation procedure.", "texts": [ " Then, we interpolate these data points to find an approximated mapping function (W,H) = f\u0303(L,D). Specific VT Shape The FEM is used to simulate tissue deformation and to predict the suture shape: (L,D) = g(W,H). The FEM is used because the suture Authorized licensed use limited to: Carleton University. Downloaded on June 04,2021 at 20:09:41 UTC from IEEE Xplore. Restrictions apply. shape is determined by the tissue deformation, which can be calculated using FEM. In the simulation, a 2-D plane-stress model is used (see Fig. 7). When the suture is generated by the VT, the suture path lies on the vertical 2-D needle insertion plane, which is on the middle side plate [see Fig. 4(a)]. Around the needle insertion plane, three side plates identically deform the cornea, and the deformed cornea is symmetrical to the needle insertion plane. Shear stress perpendicular to the needle insertion plane can be ignored because the eye fluid on the corneal surface acts as a lubricant, and the corneal surface has a small coefficient of friction [24], [25]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure9.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure9.6-1.png", "caption": "Figure 9.6 Cross-slide and top slide", "texts": [ " Adjustment for wear is provided by a tapered gib strip, which can be pushed further into the slide and slideway by the screw as wear takes place. Attached to the underside of the cross-slide is the leadscrew nut through which movement is transmitted from the leadscrew. Power feed is available to the cross-slide. The top surface contains a radial tee slot into which two tee bolts are fitted. The central spigot locates the slideway for the top slide, which can be rotated and clamped at any angle by means of the tee bolts. Graduations are provided for this purpose, Fig.\u00a09.6. Removal of equipment is carried out by rotating each cam anticlockwise until the index lines coincide and then pulling the equipment away from the spindle nose. The gearbox, fitted on the lower side of the headstock, provides the range of feeds to the saddle and cross-slide through the feed shaft, and the screw-cutting range through the leadscrew. By selecting the appropriate combination of lever positions in accordance with a table on the machine, a wide range of feed rates and thread pitches can be obtained", " Accurate movement is thus maintained relative to the centre line of the spindle and tailstock for the complete length of the bed. The top surface contains the dovetail slideway into which the cross-slide is located and the crossslide leadscrew, complete with handwheel and graduated dial, Fig.\u00a09.5. Cross-slide leadscrew and nut Figure 9.5 Saddle On the lathe shown, external dovetails are provided along each side of the cross-slide, for quick accurate attachment of rear-mounting accessories. The top slide shown in Fig.\u00a09.6, often referred\u00a0to\u00a0as the compound slide, fits on its slideway and can be adjusted for wear by means of a gib strip and adjusting screws. Movement is transmitted by the leadscrew through a nut on the slideway. A toolpost, usually four-way hand-indexing, is located on the top surface and can be locked in the desired position by the locking handle. Movement of this slide is usually quite short, 92 mm on the machine illustrated, and only hand feed is available. Used in conjunction with the swivel base, it is used to turn short tapers" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003619_tie.2021.3080207-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003619_tie.2021.3080207-Figure12-1.png", "caption": "Fig. 12. Modelling of DTPSPM machine. (a) 3D-FE model. (b) Mesh.", "texts": [ " As shown in Fig. 11(a), the start rotor positions causing the largest 3PSC peak currents stay almost unchanged when asymmetry level increases. In addition, the absolute difference between the largest and smallest 3PSC peak currents increases with asymmetry level, as shown in Fig. 11 (b). IV. FE AND EXPERIMENTAL VALIDATION A. FE Validation In order to verify the above analysis, three-dimensional FE (3D-FE) simulation is performed on a 10-pole/12-slot DTPSPM machine. The 3D-FE model is shown in Fig. 12. The parameters are listed in TABLE V and the B-H curve of laminated steel sheets is shown in Fig. 13. Firstly, the influence of start rotor position on 3PSC peak currents for the symmetrical winding configuration shown in Fig. 1(a) is simulated. Both linear and nonlinear core materials are considered. It is assumed that the initial current of faulty winding and the healthy winding current are 0A whilst the rotor speed is 5000rpm during 3PSC. Firstly, the spectra of open-circuit flux linkage of single winding group are simulated, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003340_tia.2021.3058200-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003340_tia.2021.3058200-Figure1-1.png", "caption": "Fig. 1 Machine topologies. (a) 18/17-pole IFC-BFPMM. (b) 18/13-pole IFCDSPMM.", "texts": [ " Furthermore, more experimental results such as cogging torque, static torque, and on-load torque are included in this paper. This paper is organized as follows. In Section II, machine topologies and stator/rotor-pole number combinations of both machines are discussed in detail, followed by their electromagnetic performance comparison in Section III, including flux linkage, back-EMF, inductances, and so on. In Section IV, an 18/13-pole IFC-DSPMM is manufactured and tested to experimentally validate the predicted results. Finally, some conclusions are drawn in Section V. Fig. 1 shows the topologies of both machines. According to [15], the stator and rotor poles number of IFC-DSPMM are set as 18 and 13 to suppress torque ripple. For fair comparisons, the stator pole number of IFC-BFPMM is also chosen to be 18. For 3-phase 18-stator-pole IFC-BFPMMs, when the number of rotor poles is near the number of stator poles, IFC-BFPMMs have better torque performance [11]. Fig. 2 shows the influence of the rotor pole on torque characteristics. When the number of rotor poles is 17, since IFC-BFPMM has the minimum torque ripple and relatively large average torque, the 18/17- stator/rotor-pole number combination is adopted to IFCBFPMM", " > REPLACE THIS LINE WITH YOUR PAPER IDENTIFICATION NUMBER (DOUBLE-CLICK HERE TO EDIT) < 3 stator slot, thus potentially boost output torque. Meanwhile, this leads to the unequal area of stator slots with and without PMs in 18/13-pole IFC-DSPMM. Poor space utilization suffers from unbalanced stator slots. The way to overcome this deficiency is that all side teeth bodies are bent to be parallel to middle teeth. Adjusting adequately the relative position between side teeth bodies and middle teeth can make each stator slot equal. Adjacent two coil positions in mechanical degrees m, as illustrated in Fig. 1, can be expressed as m=360/Ns (1) where Ns is the number of stator pole. The adjacent two coil positions in electrical degrees e can be deduced from the mechanical degreesm as e=Nrm (2) where Nr is the number of rotor pole. According to (1) and (2), the coil back-EMF phasor distributions of both machines are shown in Fig. 3. Coil back-EMF phasors n and n\u2019 have the opposite polarity. In 18/17-pole IFC-BFPMM, the reverse connections occur in coils exhibiting a phase shift of 180 elec. deg. (e.g. coils 1 and 10), which succeeds in eliminating even order phase back-EMF harmonics and the resultant phase backEMF is more sinusoidal", " 8 compares the open-circuit phase flux linkages between two machines with the same number of turns per coil. Both machines have the bipolar and sinusoidal phase flux linkages since the coils differing by 180 elec. deg. are connected reversely in each phase, eliminating the DC and even harmonic components. Meanwhile, as shown in Fig. 8(b), the 18/13-pole IFC-DSPMM also has a balanced three phase flux linkages. This is mainly because all coils wounded on the middle teeth of each E-shaped stator core are evenly allocated to three phases in such a machine [see Fig. 1(b) and Fig. 3(b)]. In addition, the 18/13-pole IFC-DSPMM has 63% higher fundamental component than the 18/17-pole IFC-BFPMM. Fig. 9 shows the open-circuit phase back-EMF waveforms of both machines at rated 500 r/min. As shown in Fig. 9(a), the phase back-EMF waveforms of the two machines are close to sinusoidal, which is evidenced by the very small total harmonics distortion (THD) of both machines [see Fig. 9(b)]. It should be noted that the phase back-EMF is not an ideal sinusoidal waveform due to the existence of odd harmonics" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001507_iccas.2015.7364708-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001507_iccas.2015.7364708-Figure2-1.png", "caption": "Fig. 2: Experimental setup: (a) Impact hammer, (b) and (c) acceleration sensor attached in the axial direction and the radial direction, respectively", "texts": [ " 1 shows the components of the CSF-20 model. The wave generator includes an el liptical shaped-bearing plug and a ball bearing, and func tions as an input driver. When the bearing plug rotates around the circular spline, the elastic deformation of the flex spline occurs and the reduction gear obtains the gear reduction ratio by engaging the additional two teeth. Hence, it is essential to analyse dynamic characteristics of the flex spline in order to validate the reduction mechanism. 2.2 Modal testing Fig. 2 presents the experimental setup to measure the dynamic characteristics of the flex splines of KCSF-20 and CSF-20. The natural frequencies of both models were measured on soft ground (a piece of sponge) in free-free mode and the measured frequencies were compared with each other. Fig. 3 shows the comparison of the measured natural frequencies of the flex splines of KCSF-20 and CSF-20 along the axial direction. The natural frequency of the KCSF-20 is 1.23 times greater than that of the CSF-20. Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002715_j.polymer.2020.122973-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002715_j.polymer.2020.122973-Figure3-1.png", "caption": "Fig. 3. The test specimens used for the fatigue measurements: a) Dumbbell, b) Concave.", "texts": [ " In order to compare the lifetime of both geometries, the load level of 100 \u00b1 100 N was selected. Here the peak to peak amplitude is 200 N. For the dumbbell, the peak to peak amplitude was increased to 300 N and 400 N, since the reduction of the peak to peak amplitude leads to very long measurement times. For the concave, the largest peak to peak amplitude was used at 200 N. Increasing the amplitude for the concave sample was not possible for the investigated material, because higher amplitudes exceed the permissible piston travel of \u00b1100 mm. Fig. 3 shows the dumbbell and the concave sample. Both test samples are rotationally symmetrical. The dumbbell has a transition area with a radius of 5 mm. The concave has two transition areas, the first transition area is a quarter circle with a radius of 5 mm and the second transition area is a quarter ellipse with the small half-axis of 5 mm and large halfaxis of 14.3 mm. The brackets are installed in the transition area with the radius of 5 mm by both samples. Due to the elliptical transition area by the concave sample, there will be no stress concentration near the transition area" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003261_s10514-020-09959-0-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003261_s10514-020-09959-0-Figure1-1.png", "caption": "Fig. 1 a A wheel constrained to move along a curve with its normal parallel to the center-line of the circle, i.e. their tangents are parallel and coincide. b Top-down view of the configuration in (a). c Top down view of a wheel constrained to move along a curve with its normal nonparallel to the center-line of the circle, i.e. their tangents are in different orientations", "texts": [ " Additionally, in this work, standard odometry performance tests are performed on different surfaces in order to justify the improvements of the proposed model to the kinematic performance in a systematic way. And then, the proposed model was similarly verified using another skid-steered platform with a different geometry. The proposed model contains the robot geometry as a parameter, and hence the only way to verify it is to experiment with different platforms, which is done in this paper. In order to facilitate the derivation of the kinematic model for a skid-steered wheeled platform, it is appropriate to first discuss the motion of a single wheel constrained to move along a curve (Fig. 1). Assume that a wheel of radius r and rotational speed of \u03a9 is connected with an orthogonal rigid link (an axis) of length R to a fixed point (Fig. 1a, b). Also assume that the wheel is experiencing pure rotational motion with a surface speed of u = \u03a9r , and hence its angular speed with respect to the fixed point is given by \u03c9 = u R . In this motion, the tangent of the wheel and the tangent of the curve are parallel, allowing the full displacement of the wheel perimeter to be transferred to displacements on the circle. Although two wheel differential drive robots are not pinned to a point like this wheel when they are rotating, they face the same phenomena. Skid-steered wheeled robots face a slightly different configuration. For them the link is not perpendicular to the wheel plane (Fig. 1c), and hence if the wheel is forced to follow a curve, the velocity vector of the wheel and the tangent of the curve do not align, preventing the transfer of the full displacement along the perimeter to the the curve. However, rotation of the wheel will still cause a displacement along the curve. Infinitesimally, this displacement will be at most equal to the size of the projection of the displacement of the wheel perimeter to to the curve\u2019s tangent. The size of the projection can be calculated as udt cos\u03b1 where \u03b1 is the angle between the velocity vector and the tangent of the curve, and dt is the infinitesimal time step" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.12-1.png", "caption": "Fig. 3.12 Basic Euler motions of the earth. Intrinsic rotation (green \u201cR\u201d ), precession (blue \u201cP\u201d), and nutation (red \u201cN\u201d). (From Wikipedia, original design by Dr. H. Sulzer)", "texts": [ "16) and matrix multiplication, we get RHelm = \u23a1 \u23a3 cos \u03b8H cos\u03c6H \u2212 sin \u03b8H cos\u03c6H cos\u03c8H + sin \u03c6H sin\u03c8H sin \u03b8H cos\u03c6H sin\u03c8H + sin \u03c6H cos\u03c8H sin \u03b8H cos \u03b8H cos\u03c8H \u2212 cos \u03b8H sin\u03c8H \u2212 cos \u03b8H sin \u03c6H sin \u03b8H sin \u03c6H cos\u03c8H + cos\u03c6H sin\u03c8H \u2212 sin \u03b8H sin \u03c6H sin\u03c8H + cos\u03c6H cos\u03c8H \u23a4 \u23a6 (3.27) When R is given, the Helmholtz angles (\u03b8H , \u03c6H , \u03c8H ) can be using \u03b8H = arcsin(Ryx ) \u03c6H = \u2212 arcsin ( Rzx cos \u03b8H ) \u03c8H = \u2212 arcsin ( Ryz cos \u03b8H ) . (3.28) Yet another sequence to describe 3-D orientation is common in theoretical physics and mechanics and in other technical literature, and often referred to as Euler sequence.5 In order to describe the movement of a spinning top rotating on a table, or of the earth during its rotation around the sun (see Fig. 3.12), three angles are needed: the intrinsic rotation (\u03b3 ), nutation (\u03b2), and precession (\u03b1). Using these three angles, the orientation of the spinning object is described by (see Fig. 3.13) \u2022 a rotation about the z-axis, by an angle \u03b1, \u2022 followed by a rotation about the rotated x-axis, by an angle \u03b2, and \u2022 followed by a rotation about the twice-rotated z-axis, by an angle \u03b3 . REuler = Rz(\u03b1) \u00b7 Rx (\u03b2) \u00b7 Rz(\u03b3 ). (3.29) This leads to the parametrization REuler = \u23a1 \u23a3 \u2212 sin \u03b1E cos\u03b2E sin \u03b3E + cos\u03b1E cos \u03b3E \u2212 sin \u03b1E cos\u03b2E cos \u03b3E \u2212 cos\u03b1E sin \u03b3E sin \u03b1E sin \u03b2E sin \u03b1E cos \u03b3E + cos\u03b1E cos\u03b2E sin \u03b3E \u2212 sin \u03b1E sin \u03b3E + cos\u03b1E cos\u03b2E cos \u03b3E \u2212 cos\u03b1E sin \u03b2E sin \u03b2E sin \u03b3E sin \u03b2E cos \u03b3E cos\u03b2E \u23a4 \u23a6 (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001765_1464419315584709-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001765_1464419315584709-Figure1-1.png", "caption": "Figure 1. Reduced model of ball and race contact.", "texts": [ "3,4,23 The normal contact forces between balls and races are calculated using both numerical solution of EHL and Hertz contact theory, and the other interactions between the bearing elements are computed using the same model. The results are compared to show the effects of the EHL on the dynamics of ball bearings. Dynamic characteristics of elastohydrodynamic lubricated point contacts When the vibration amplitude is small, the lubricating film between balls and races can be viewed as a linear spring\u2013damper system, as shown in Figure 1. The mutual approach can be derived from the steady state EHL solution, and the relation between mutual approach and load, entrainment velocity can be used in the dynamic analysis of ball bearings directly. The damping coefficient is calculated based on the transient solution. The governing equations of EHL theory include Reynolds equation, film thickness equation, lubricant viscosity\u2013pressure relation, density\u2013pressure relation, and motion equation. Applying the general assumptions for hydrodynamic lubrication problem, and ignoring the side leakage of the lubricant in the y-direction, the Reynolds equation used in this paper reads @ @x h3 @p @x \u00fe @ @y h3 @p @y \u00bc 12u @ h\u00f0 \u00de @x \u00fe 12 @ h\u00f0 \u00de @t \u00f01\u00de where is the oil density, p is oil film pressure, h is oil film thickness, is oil viscosity, and u is entrainment velocity" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002862_tase.2020.3031691-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002862_tase.2020.3031691-Figure2-1.png", "caption": "Fig. 2. Manipulator of the concrete pump with virtual joints ( ) and vector from the frame i to the successor frame ( ).", "texts": [ " For example, the Jacobian matrix Jw i v is denoted by the right subscript v to indicate a translational Jacobian, the right superscript i to specify the affiliation with the frame i , and the left superscript w to denote the frame of resolution of the components. A table of notation can be found in Appendix C. The manipulator of the concrete pump comprises five segments that are connected by single degree of freedom revolute joints \u03b8 i , i = 1, . . . , 5 about the horizontal z-axis of the local right-handed Cartesian frame i , as shown in Fig. 2. An additional revolute joint \u03b80 at the fixed base allows rotating the boom around the vertical y-axis of frame 0. Each horizontal joint is actuated by a double acting hydraulic cylinder, which is connected through a mechanical linkage. The relative cylinder displacement between housing and rod is denoted by si , i = 1, . . . , 5. The vertical joint at the base has a hydraulic rotary motor with a gearbox. All joint variables are collected into the vector \u03b8 \u2208 R 6 and named the actuated coordinates", " They are named the elastic coordinates. The resulting lumped-parameter model approximates deformation in the vertical plane. The complete derivation of the dynamic model of the concrete pump can be found in [40]. In contrast to [40], the virtual joints are moved off-center in the local y-axis direction to more closely represent the shape of the links. The x-axis of the frames i is aligned parallel to main axial direction of the links. The inertial frame I matches the frame 0 for \u03b80 = 0 (its position is shifted in Fig. 2 for clarity). The actuated and elastic coordinates form the vector of generalized coordinates q = [\u03b8T \u03b4T ]T \u2208 R 11. The differential forward kinematics of the frame i yields vi = J i v (q)q\u0307 (1a) \u03c9i = J i \u03c9(q)q\u0307, i = {0, . . . , 5, tcp} (1b) where vi , \u03c9i \u2208 R 3 are the translational and rotational velocities of frame i defined with respect to the frame I and J i v , J i \u03c9 are the corresponding translational and rotational Jacobians, respectively. Typically, inverse differential algorithms assume a fully actuated rigid body system, i", " The time derivative of (4) yields s\u0307i = ki k(\u03b8 i )\u03b8\u0307 i (5) with ki k = (\u2202pi/\u2202\u03b8 i ). Equation (5) establishes a linear relationship between the rotational joint velocity \u03b8\u0307 i and the translational cylinder velocity s\u0307i . For efficiency, the relations (4) and (5) are discretized and used with a linear interpolation scheme. To simplify the graphic representation in the following, the manipulator is plotted as line segments between the actuated joints i , such as in Fig. 4. The line segments are also depicted as dashed lines in Fig. 2. Although the lines Authorized licensed use limited to: Rutgers University. Downloaded on May 18,2021 at 05:07:28 UTC from IEEE Xplore. Restrictions apply. are straight, they describe the configuration-dependant link deformation by varying positions of the successor frame. The cQP of the TCP control algorithm is defined as min \u03b8\u0307 f1(\u03b8\u0307) + f2(\u03b8\u0307) + f3(\u03b8\u0307) (6a) f1(\u03b8\u0307) = \u2016v\u0303e \u2212 Je\u03b8\u0307\u2016W v\u0303 (6b) f2(\u03b8\u0307) = \u2016\u02d9\u0303 \u03b8 c \u2212 J c\u03b8\u0307\u2016W \u03b8\u0303 (6c) f3(\u03b8\u0307) = \u2016\u03b8\u0307\u2016W \u03b8 (6d) subject to \u03b8\u0307 \u2264 \u03b8\u0307 \u2264 \u03b8\u0307 (6e) C \u03b8\u0307 \u2264 Q (6f) where \u03b8\u0307 is the optimization variable, v\u0303e \u2208 R ne is the desired operation space velocity, J e \u2208 R ne\u00d76 is the operation space Jacobian, \u02d9\u0303 \u03b8 c \u2208 R 6 are the desired joint velocities, Jc \u2208 R 6\u00d76 is the configuration control Jacobian, W v\u0303 \u2208 R ne\u00d7ne and W \u03b8\u0303 \u2208 R 6\u00d76 are positive semidefinite weighing matrices, W\u03b8 \u2208 R 6\u00d76 is a positive definite weighing matrix, \u03b8\u0307 , \u03b8\u0307 \u2208 R 6 are the lower and upper joint velocity bounds, respectively, and C \u2208 R 64\u00d76 is the hydraulic flow constraint matrix with the upper bound Q \u2208 R 64" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001762_j.ifacol.2016.03.134-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001762_j.ifacol.2016.03.134-Figure2-1.png", "caption": "Fig. 2. Twin rotor MIMO system laboratory model (TRMS (1997))", "texts": [ " Finally, using the combination of first level estimates of plant parameter vector \u03b8\u0302k with adaptive weights W (t) at every instant, the virtual estimate of plant parameter values \u03b8s can be obtained as in (20) and can be used to find the new control input as us = As(x, \u03b8s) \u22121[\u2212Bs(x, \u03b8s) + v\u0302j ] (30) In the above equation, As(x, \u03b8s) is assumed to be bounded away from zero, because of the fact that \u03b8s always resides inside the convex hull of \u03b8\u0302k (Sastry and Isidori (1989) Sastry and Bodson (1989)). As discussed in comment (c) of theorem 7.3.1 in (Sastry and Bodson (1989)), the parameters \u03b8\u0302k are kept in a certain range, such that, at every point inside their convex hull, q(x, \u03b8s) is bounded away from zero. The basic control block diagram with second level adaptation is shown in Fig. 1. A TRMS as shown in Fig. 2 is simulated to validate the theoretical developments discussed in previous sections. A TRMS has two rotors, main rotor and a tail rotor. The aim here is to control the attitudes in two directions i.e. pitch and yaw control. The dynamical equation of TRMS in state space form is given as (TRMS (1997)): o he o . x\u03071 = x2 x\u03072 = a1 I1 x5 2 + b1 I1 x5 \u2212 Mg I1 sinx1 \u2212 B1\u03b1v I1 x2 + 0.0326 2I1 sin(2x2)x4 2 \u2212 kgy I1 a1 cos(x1)x4x5 2 \u2212 kgy I1 b1 cos(x1)x4x5 x\u03073 = x4 x\u03074 = a2 I2 x6 2 + b2 I2 x6 \u2212 B1\u03b1h I2 x4 \u2212 kca1 I2 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003001_j.engfailanal.2020.105195-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003001_j.engfailanal.2020.105195-Figure13-1.png", "caption": "Fig. 13. Maximal principal stress distribution map in the bicycle frame for: a) pedalling force F1 (angle q = 7.5\u25e6), b) horizontal force F2, c) vertical force F4, unit in MPa.", "texts": [ " The calculations were carried out in the structural analysis module at constant loading and reaction conditions. The range of analyses includes the identification of stress distribution for boundary conditions specified in Section 4.1. The stresses in the element were calculated in the integration points, extrapolated to the nodes and indicated as mean values. Fig. 10 shows the distribution map of the maximum main stresses. The position of maximum values (S(1), S(2), S(3)) was indicated for all loading cases. Fig. 13a shows the stress distribution map for loading applied to the pedal. The position of the maximum stress corresponds to the location of fatigue failure (Fig. 3). The stress was calculated for selected angle values (q = 7.5; 15; 22.5; 30\u25e6). The value of a maximum nominal stress increases with an increase in angle q. T. Tomaszewski Engineering Failure Analysis 122 (2021) 105195 Table 4 shows the maximum stress for all loading cases, relative stress (Smax/Re) and recommended fatigue life. In other loading cases (Fig. 13b, c), the maximum stress was significantly below the material\u2019s yield strength. Fig. 14 shows the local stress distribution taking into account the geometry of the welded joint at the fatigue crack location. The boundary conditions are consistent with the first case for the force F1 applied to the pedal axis, directed downwards at an angle 7.5\u25e6 (Fig. 12a). The stress distribution along the tube is schematically shown below the distribution map. The maximum stress is located on the inside of the top tube" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000342_j.triboint.2019.105881-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000342_j.triboint.2019.105881-Figure5-1.png", "caption": "Fig. 5. Hot oil carry-over.", "texts": [ " Other equations, such as the lubricant temperature\u2013viscosity equation, the thermoelastic deformation equation of the pad, see Ref. [18]. 3D-TEHD calculation is multi-field coupled calculation, and the thermal influence of oil film and the deformation of bearing bush are considered, so it is a more comprehensive bearing calculation model. One of the most important boundary conditions of 3D-TEHD analysis is the oil film inlet temperature boundary condition. For the supply oil temperature boundary condition, Formula (9-1) becomes =T Ti s (10) For the hot oil carry-over model, as shown in Fig. 5, Formula (9-1) is: \u239c \u239f= + \u239b \u239d \u2212 \u239e \u23a0 \u22c5 \u2212T K Q Q TT 1 ( T )s q l i s oi (11) The carry-over coefficient Kqis used to represent the hot oil carried into the downstream pad, where. = \u2212 \u22c5 \u2212Q Q K Q Q( )s i q i l Another hot oil carry-over coefficient K is easier to calculate: = \u2212 \u2212 = +K T T T T T T T( )/( ), 1 2 ( )i s r s r i o So = \u2212 \u2212 + \u2212 T K K T K K T2(1 ) 2 2i s o (12) In practice, this can be simplified as: = + \u22c5 \u2212T T \u03b2 T T( )i s o s (13) Where = \u2212( )\u03b2 K K2 , \u03b2varies from 0 to 1. For the thermal boundary layer inlet temperature, assuming that the temperature of the runner thermal boundary layer is distributed along the depth direction as a third-order polynomial, and according to the integral equation of the boundary layer of the plate, the thickness of the thermal boundary layer can be given as: = \u22c5 \u22c5 \u22c5 \u2212 \u2212 \u03b4 x R P4" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure11.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure11.7-1.png", "caption": "Figure 11.7 Milling-machine controls", "texts": [ " These are capable of sliding and swinging aside to provide access in order to load/unload the workpiece, change cutters or remove swarf. The high-impact polycarbonate panels provide all-round visibility. Electrical safety interlocks are available which prevent the machine being started if the guard is in the open position and stop the machine should the guard be opened when the machine is running. Figure 11.6 Slide and swing-aside guard The various controls of a typical horizontal milling machine are shown in Fig.\u00a011.7. These are identical to those of a vertical machine. Spindle speeds are selected through the levers (4), and the speed is indicated on the change dial (5). The speeds must not be changed while the machine is running. An \u2018inching \u2019 button (3)\u00a0is situated below the gear-change panel and, if depressed, \u2018inches\u2019 the spindle and enables the gears to slide into place when a speed change is being carried out. Alongside the \u2018inching\u2019 button is the switch for controlling the cutting-fluid pump (1) and one for controlling the direction of spindle rotation (2)", " A range of sizes is available from This cutter has helical teeth on the circumference and teeth on one end and is used for light operations such as milling slots, profiling and facing narrow surfaces, Fig.\u00a011.8(i). The end teeth on non-centre-cutting cutters are not cut to the centre, so this cutter cannot be fed in a direction along its own axis. Centre-cutting types have teeth cut to the centre which allows drilling and plunging operations. The switch panel, situated on the front of the knee, contains a black button (B) to start the feed motor, Fig.\u00a011.7. This is provided to facilitate setting up when feed movements are required without spindle rotation. The green button (G) starts the spindle and feed motors, while the mushroom-headed red button (R) provides the means of stopping the machine. There are many different types of milling cutters available, and for convenience they can be classified according to the method of mounting: those with a central hole for mounting on an arbor, those with a screwed shank for holding in a special chuck and the large facing cutters which mount directly on to the spindle nose" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001137_ipec.2014.6869708-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001137_ipec.2014.6869708-Figure6-1.png", "caption": "Fig. 6. Suspending force generation principle of proposed BLSRMs.", "texts": [ " 5, mathematics model is not used in the control scheme, only a PI type speed controller is adopted to regulate the motor speed, and two independent c1ose loop PID air-gap displacement controllers, one for x direction and the other for y-direction, are used to generate the desired suspending force commands Fx * and Fy * to keep the rotor at the center position. Further, in the control scheme, the actual current values of the exciting phase of the suspending force winding can be controlled through the hysteresis method according to the command current signals. Fig. 6 shows the suspending force generation principle of the proposed BLSRMs. As shown in Fig. 6, if the winding on the suspending force pole Pxp is excited, the suspending force F:, in the positive x-direction can be generated. If the winding on the suspending force pole Pyp is excited, the suspending force Fy in the positive y direction can be generated. In the same way, if the windings on the suspending force pole P,p and Pyp are excited simultaneously, the force F will be generated, which is the synthesis force of Fx and Fy - Furthermore, the value and direction of the force F can be regulated by changing the values of the currents in the two windings", " Therefore, when controlling values of the currents in the four suspending force windings, the desired resultant force in any arbitrary direction and magnitude can be obtained to compensate for the unbalanced pull force caused by the non-uniform air-gap. According to the suspending force generation principle in the proposed BLSRMs, the force F, in any direction and magnitude, can be generated by the windings on the two poles, i.e. one each in the x- and y-directions. For instance, assume that F is the desired force, as shown in Fig. 6. In this case, because F is in the first quadrant, pole Pxp and Pyp are selected to produce the suspending force. Table I shows the excitation poles for various forces. In the table, F, and Fy are used to determine the suspending force control poles. After choosing the force control poles, the command current ix * and iy * for these two poles can be calculated by a 3-dimensional look-up table. According to the above suspending force current calculations, the asymmetric converter based on hysteresis control method is applied to control the winding current" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001821_s0263574716000345-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001821_s0263574716000345-Figure2-1.png", "caption": "Fig. 2. A parallelepiped spanned by three weighted screws.", "texts": [ " Rank ( MTQM ) = Rank ( MTM ) = Rank (M) (1) Hence, the weighted twists or weighted wrenches can be adopted in the screw matrix to detect singularities of a screw system. In the remainder of this paper, screws used in dot product or determinant are all represented as the weighted screws. Weighted screws in an N order system span an N dimensional polyhedron whose edges are the N screw vectors, such as a parallelogram for a two order system in Fig. 1 and a parallelepiped for a three order system in Fig. 2. It can be proved that the volume of the N dimensional polyhedron, Vpol , can be expressed as follows. Vpol = \u221a det ( MTM ) , (2) where M denotes the weighted screw matrix. Another N dimensional rectangular polyhedron is established with its edges along coordinate axes, and the length of each edge is expressed as \u2016$i\u2016, i = 1, 2, . . . , N . The volume of this rectangular http://dx.doi.org/10.1017/S0263574716000345 Downloaded from http:/www.cambridge.org/core. New York University Libraries, on 10 Dec 2016 at 01:23:18, subject to the Cambridge Core terms of use, available at http:/www", " In order to compare the closeness to singularities of screw systems with different orders, the geometric average of the normalized volume should be adopted as the metric, which is defined as follows. \u03be = (I+J\u22121) \u221a \u03c7 = \u239b \u239c\u239c\u239c\u239d \u221a det ( MTM ) I\u220f i=1 \u2225\u2225$i \u2225\u2225 \u00b7 J\u220f j=1 \u221a det ( MT j Mj ) \u239e \u239f\u239f\u239f\u23a0 1 I+J\u22121 . (14) The symbols in Eq. (14) are coincident with the symbols in Eqs. (6), and (14) is the (I + J -1)th root of Eq. (6). Note that the power in Eq. (14) is 1/(I+J-1), not 1/(I+J), since this geometric average value can be related to the average characteristic angle between screws. In Fig. 1, there is an angle \u03b1 for two order screw system; in Fig. 2, there are two angles \u03b1 and \u03b2 for three order screw system. For the manifold spanned by I definite screws, there are I -1 characteristic angles between these screws, as shown in Figs. 1 and 2. For the manifold spanned by J submanifolds whose base screws are indefinite, there are J -1 characteristic angles between these submanifolds, but there are no characteristic angles within each submanifold since its base screws are indefinite. Hence, for the entire manifold consisting of I definite screws and J submanifolds, there are I + J -1 characteristic angles between these screws and submanifolds" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003249_tte.2021.3049466-Figure18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003249_tte.2021.3049466-Figure18-1.png", "caption": "Fig. 18. Prototype and test rig. (a) Dual stators. (b) Rotor. (c) Test rig. \u2460: general inverter. \u2461: drive-load platform. \u2462: inverter. \u2463: control unit.", "texts": [ " In brief, with the schemes proposed in this article, the HS-RDPM can have higher torque density than conventional FTPMs and ideal performances of fault tolerance. The performance comparisons are listed in Table V. Authorized licensed use limited to: Dedan Kimathi University of Technology. Downloaded on May 18,2021 at 11:03:36 UTC from IEEE Xplore. Restrictions apply. To verify the simulations, a prototype HS-RDPM has been designed and manufactured, and its test rig has been built. The laminations of dual stators and rotor frame are shown in Fig. 18(a) and (b). The stator composes of the inner stator and outer stator, both of which adopt the hybrid slot. The test rig, shown in Fig. 18(c), mainly consists of a general inverter, drive-load platform, inverter, and control unit. The drive-load platform contains the HS-RDPM and an induction machine. The inverter receives the order of the control unit and drives HS-RDPM. The general inverter drives the induction machine that acts as the load of HS-RDPM. Fig. 19 shows the measured open-circuit back EMF under 1000 rpm, whose amplitude is about 14.6 V and slightly smaller than the simulated value, about 15.1 V in Fig. 9(a). Fig. 20 shows the simulated and measured average torques under different currents, and it can be drawn that the measured values are very close to the simulated values" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001937_j.optlastec.2016.08.011-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001937_j.optlastec.2016.08.011-Figure2-1.png", "caption": "Fig. 2. Initial meshing of metal droplet for FEM analysis.", "texts": [ " The three-dimensional finite element model developed in this paper is based on the minimum energy principle. The finite element calculations are performed using Surface Evolver, an interactive program for the study of liquid surfaces shaped by various energies, including surface tension energy, gravitational energy and many others [30,31]. The gravitational potential energy of a droplet Eg can be given by \u222e \u03c1= ( )E gzdV 1g V m where \u03c1m is the density of liquid metal, g is the gravitational acceleration, z is the distance between the droplet and the zero potential energy surface (as shown in Fig. 2) and V is the volume of the droplet. According to Gauss Formula, Eg can also be expressed by surface integral as follows \u222e \u03c1= \u2192 \u2192 ( )E gz k dS 1 2 2g S m 2 where S is the surface of the droplet and \u2192 k is the unit vector along the z axis. The surface potential energy \u03b3E generated by surface tension is given by \u222e \u03b3= \u22c5 ( )\u03b3E dS 3s where \u03b3 is the surface tension coefficient. ) MIG welding, and (b) laser-MIG hybrid welding. In laser-MIG hybrid welding, plasma drag force Fp promotes droplet transfer, and it can be given by [29] \u03c1 = ( ) \u239b \u239d \u239c\u239c \u239e \u23a0 \u239f\u239fF C A v 2 4 P D P f f 2 where CD is the Reynold number, \u03c1f is the density of plasma fluid, vf is the velocity of plasma fluid and AP is the action area of plasma fluid which can be given by [29] ( )\u03c0= \u2212 ( )A R R 5P d w 2 2 where Rd is the radius of the droplet and Rw is the radius of the necking part of the droplet", " The potential energy generated by plasma drag force Ep can be calculated by \u222e \u03c1= ( + ) ( )E a y a z dV 6p V py pz m where apy and apz are the components of droplet acceleration generated by plasma drag force along the y and z axis, respectively. Then, according to Gauss Formula and Eq. (4) and (5), Ep can be eventually calculated by ( )\u222e \u03c1 \u03b8 \u03b8= \u2212 \u22c5 \u2192 + \u22c5 \u2192 \u22c5 \u2192 ( ) \u239b \u239d\u239c \u239e \u23a0\u239fE C R R v R y j z k dS 3 8 sin 2 cos 2 7 p S D d w f f d 2 2 2 3 2 2 where \u03b8 is the angle between wire axis and z axis (as shown in Fig. 2) and \u2192 j is the unit vector along the y axis. The vaporization-induced recoil force FRL can be calculated by [29,32,33] ( ) ( ) ( ) ( ) ( ) \u03c0 \u03c1 = \u2212 \u00d7 \u2212 \u2264 > \u23a7 \u23a8 \u23aa\u23aa\u23aa \u23a9 \u23aa\u23aa\u23aa \u239b \u239d \u239c\u239c \u239e \u23a0 \u239f\u239f 8 F R C A V N k T M B M L N k T D R D R D R 1 4 exp / exp /2 0 RL h D m a B s a a v a B s LA h LA h LA h 2 2 0 2 3/2 0 2 2 where Rh is the largest radial dimension of metal vapor, A is the area of droplet projection image which is vertical to flow direction, V0 is a constant whose value is of the order of the speed of sound in the condensed phase, Na is Avogadro\u2019s number, kB is Boltzmann\u2019s constant, Ts is the melt's surface temperature, Ma is the molecular weight of metal vapor, B0 is a vaporization constant, Lv is the latent heat of evaporation and DLA is the laser-wire distance", " Similar to the potential energy generated by plasma drag force, the potential energy Em generated by hybrid electromagnetic force can be calculated by \u222e ( ) ( ) ( ) \u03c0 \u03b8 \u03b1 \u03b8 \u03b1 = \u2212 + + \u2212 \u22c5 \u2212 \u2212 \u00d7 + \u2192 + + \u2192 \u22c5 \u2192 ( ) \u239b \u239d\u239c \u239e \u23a0\u239f E I P I P I P R y j z k dS 152.6308 2.0481 128.4058 0.4618 0.0011 5.196 3 4 sin 2 cos 2 12 m S d 2 2 3 .2 2 In order to add critical additional axial acceleration to the model, it is assumed that critical additional axial acceleration is equal to n times the gravitational acceleration, that is, = \u22c5a n gc . Similarly, the potential energy Ec generated by additional mechanical forces can be calculated by \u222e \u03c1 \u03b8 \u03b8= \u2192 + \u2192 \u22c5 \u2192 ( ) \u239b \u239d\u239c \u239e \u23a0\u239fE n g y j z k dSsin 2 cos 2 13 c S m 2 2 Fig. 2 shows the initial meshing of the droplet in hybrid welding for finite element modelling analysis. For the convenience of meshing, the three-dimensional coordinate system is rotated \u03b8 degrees around X axis and the wire axis is the Z0 axis of the new coordinate system. The initial mesh consists of nodes (V1\u2013V12), edges (E1\u2013E20) and surfaces (F1\u2013F11). Boundary conditions for the welding wire and droplet are discussed as follows. The welding wire is a cylinder, so the shape constraint for the wire is given as \u2032 + \u2032 = ( )x y R 14w 2 2 2 where Rw is the radius of the wire" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure14-1.png", "caption": "Fig. 14. Evaluation points of dynamic response.", "texts": [ " In the calculation, the input speed is 1500 r/min, the input power is 45 kW, and the eccentricity errors of gear 1 are 0 and 20 \u03bcm. Dynamic response of the gearbox is calculated with the mode superposition method in ANSYS software. In the calculation, the damping ratio of structure is 0.05, the time step is \u2206t = 0.00025 s, the total number of steps is 2048, and the total solution time is 0.512 s. The first 200 orders are taken for modal superposition to calculate the vibration response in the time domain. FFT is applied to vibration acceleration for obtaining the response curves in the frequency domain. Fig. 14 shows the locations of vibration response points on the housing (their positions are same as the vibration test points in the experiment). Points 1#-10# are located upon 10 bearing seats of the gearbox. The RMS values for vibration acceleration of all points with different eccentricity error values are listed in Tables 5 and 6. As shown in Tables 5 and 6, the maximum vibration acceleration of the gear box appears in Z-direction of the calculation point 6, which is located at the high-speed bearing seat", " 16, the gearboxes adopt Z-shaped installation method with two fulcrums fixed, that is, one side of the input stage closed to the input motor is fixed by a stent, and one side of the output stage closed to the test gearbox is fixed by a stent. In the process of working, the gears of both gearboxes transmit a constant level of torque with the drive motor having the sole task of overcoming the power losses of the transmission system to achieve a desired speed level. The experimental load is applied through a dynamometer. The test points of the experiment are shown in Fig. 14. The bearing plays an important role in the gearbox, and the load received by the transmission system is transferred to the gearbox through the bearing. Thus, test points are located upon 10 bearing seats of the gearbox. Before the vibration test, the measuring points shall be polished smooth and the surface oil stains shall be wiped clean. Reasonable wiring shall be conducted depending on the distribution of the measuring points without affecting the normal operation of the gearbox. Under the rated working condition (1500 r/min, 45 kW), the test run is more than 30 min to ensure that the box reaches the heat balance state" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure1-1.png", "caption": "Figure 1. The full disc", "texts": [ " With this purely mechanical analysis, we determine the variation of the total deformation of the model disc-pads at the time of braking, the stress distribution of the brake pads. Lastly, the results of the thermoelastic coupling such as Von Mises stress, Contact pressure field and the total deformations of the disc and pads are also presented in this study which will be useful for the design in the industry of automobile. The friction tracks are known as external when they are located on the side of the rim and interior when they are located on the side of the axle. The disc consists of a solid ring with two friction tracks (Fig. 1), a bowl which is fixed to the hub and one which is fixed the rim and a connection between the tracks and the bowl. This connection is necessary because the ring and bowl part which is fixed at the hub are not on the same plane for reasons of obstruction and housing of the pads and caliper. The junction between the bowl and the tracks is often machined into the shape of threat to limit the heat flux resulting from the tracks towards the bowl in order to avoid an excessive heating of the rim and the tire" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001538_iecon.2015.7392832-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001538_iecon.2015.7392832-Figure2-1.png", "caption": "Fig. 2. Meshed FEA model of the studied SynRM by JMAG", "texts": [ "T a opt ds opt b opt qs opt c opt i i i p i i \u03b8= P (13) where ( )p\u03b8P : Park\u2019s transformation is defined as cos( ) sin( ) 2 2 2 ( ) cos( ) sin( ) 3 3 3 2 2 cos( ) sin( ) 3 3 T p p p p p p p \u03b8 \u03b8 \u03c0 \u03c0 \u03b8 \u03b8 \u03b8 \u03c0 \u03c0 \u03b8 \u03b8 \u2212 = \u22c5 \u2212 \u2212 \u2212 + \u2212 + P (14) The output power outP and the efficiency of the SynRM drive \u03b7 are expressed as : .100%out out L P P P \u03b7 = + (15) III. INVESTIGATION OF ARTIFICIAL NEURAL NETWORKS FOR THE PARAMETERS ESTIMATION This section presents the investigation of Artificial Neural Networks to estimate the parameters , ,d q iL L R of the SynRM including magnetic saturation and cross-coupling. Fig. 2 shows the meshed FEA model of the studied SynRM when using JMAG software in order to calculate the parameters of the machine. The inductances ,d qL L and iron loss resistance iR obtained with FEA is shown in Fig. 3. As can be seen, when the currents dsi and qsi increase, the inductance ,d qL L will be decreased because of the magnetic saturation and cross-coupling. Fig 3c) shows the variation of the iron loss resistance following the current dsi and the rotor speed . (a) Inductance dL (b) Inductance qL (c) Iron loss resistance iR The parameters of the machine including , ,d q iL L R obtained from FEA were used to train the Neural Network is shown in the Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002905_iecon43393.2020.9254572-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002905_iecon43393.2020.9254572-Figure1-1.png", "caption": "Fig. 1: Bicycle Structure", "texts": [ " Based on which, the nonlinear Euler-Lagrange model is deduced. The simplified nonlinear descriptor model is then proposed. In Section III, with a polynomial feedback controller, the nonlinear closed-loop descriptor system is verified. The sufficient stability condition for the closed-loop system and its proof, which is the main part of this paper, are offered. In Section IV, a numerical example is provided to examine the validity of the theorem. Section V concludes the whole paper. Shown as in Fig.1, the physical representations of bicycle robot and their system states are presented in TABLE I With the theoretical proof in [1], [14], the pitch angle is neglected and treated as zero. Its effect on the model dynamics is neglected. Vector (x(t), y(t), z(t)) is the position of the bicycle robot\u2019s mass point to the defined nature ground coordinate. Vector (xg, zg) is the total mass point Ob of the bicycle robot to the coordinate original rear wheel/ground contact point. The total mass of the bicycle robot, including the mass of rear assemble mb, rear wheel mr, (steering) handlebar assemble mh and front wheel mf , is mT = mb +mf +mr +mh", " With (1), one can obtain that \u03c3 = 1 R = tan(\u03b4(t))cos(\u03bb) Wcos(\u03c8b(t)) (2) And, with the nonholonomic constraint for the rear wheel contact point, it can be obtained that{ vr = R\u03c8\u0307h = \u03c8\u0307h Wcos(\u03c8b(t)) tan(\u03b4(t))cos(\u03bb) v\u22a5 = 0 (3) \u03c8h is an intermediate variable represented the heading angle of the bicycle robot. Further, one can obtain the velocity of the mass center of the bicycle robot that [10], [15] vg = (vr + zg\u03c8\u0307hsin(\u03c8b(t)))i +(v\u22a5 + xg\u03c8\u0307h \u2212 zg\u03c8\u0307bcos(\u03c8b))j + (zg\u03c8\u0307bsin(\u03c8b))k The Euler-Lagrangian kinetic and potential energy of the bicycle robot is L = 1 2 (mT v 2 g + Js\u03b4\u0307 2)\u2212mT g(hcos(\u03c8b)\u2212\u2206hG) (4) where \u2206hG = \u03b3xgsin(\u03c8b(t)) \u2248 c\u03b4(t)cos(\u03bb) Wcos(\u03c8b(t)) xgsin(\u03c8b(t)) With W sin(\u03b4g) = c sin(\u03b3) c is the trail defined in Fig.1, one can obtain the approximation of variable \u03b3 that \u03b3 \u2248 csin(\u03b4g) W = c tan(\u03b4(t))cos(\u03bb) Wcos(\u03c8b(t)) cos(\u03b4g) \u2248 c\u03b4(t)cos(\u03bb) Wcos(\u03c8b(t)) 2793 Authorized licensed use limited to: Central Michigan University. Downloaded on May 14,2021 at 02:08:05 UTC from IEEE Xplore. Restrictions apply. With the nonholonomic constraint (3), one can obtain the approximation of \u03c8\u0307h that \u03c8\u0307h \u2248 vr \u03b4(t)cos(\u03bb) Wcos(\u03c8b(t)) and Lagrange function (4), one can obtain that LC = 1 2 (mT v 2 g + Js\u03b4\u0307 2)\u2212mT g(hsin(\u03c8b)\u2212\u2206hG) = cgmT xg cos(\u03bb)\u03b4(t) sin(\u03c8b(t)) W cos(\u03c8b(t)) \u2212mT xgzg cos(\u03bb)(\u03c8\u0307b(t)) sin(\u03b4(t))vr(t) W cos(\u03b4(t)) +1/2 (cos(\u03bb))2(vr(t)) 2mT zg 2 W 2 +1/2mT zg 2 ( \u03c8\u0307b (t) )2 + mT zg cos(\u03bb) sin(\u03b4(t)) sin(\u03c8b(t))(vr(t)) 2 W cos(\u03b4(t)) cos(\u03c8b(t)) \u2212ghmT cos (\u03c8b (t)) \u22121/2 (cos(\u03bb))2(vr(t)) 2mT (xg 2+zg 2) W 2(cos(\u03c8b(t))) 2 \u22121/2 (cos(\u03bb))2(vr(t)) 2mT zg 2 W 2(cos(\u03b4(t)))2 +1/2 Js ( \u03b4\u0307 (t) )2 + 1/2mT (vr (t)) 2 +1/2 (cos(\u03bb))2(vr(t)) 2mT (zg 2+xg 2) (cos(\u03b4(t)))2W 2(cos(\u03c8b(t))) 2 (5) One can obtain the Euler-Lagrange model M(q)\u03be\u0307 + C(\u03be(t), q(t)) +G(q) = \u03c4 (6) where \u03be(t) = [\u03c8\u0307b(t) \u03b4\u0307(t) vr(t)] T= [q\u0307T (t) vr(t)] T , q(t) = [\u03c8b(t) \u03b4(t)] T , vr is the forward velocity of the rear body, \u03c8b is the roll angle of rear body, \u03b4 is the angle of the steering axis (also named as handlebar)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001875_978-4-431-54595-8-Figure9.11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001875_978-4-431-54595-8-Figure9.11-1.png", "caption": "Fig. 9.11 Skin sensor", "texts": [ " The microphones capture human voice by forming virtual sound beams. The floor sensors that cover the entire space of the laboratory reliably detect footsteps. Skin sensors are often installed on robots. Soft and sensitive skin sensors are particularly important for interactive robots. However, there is no skin sensor to sufficiently mimic human sensitive and soft skin. We have developed novel skin 206 H. Ishiguro sensors for our androids. The sensors are made by combining silicone skin and piezo films, as shown in Fig. 9.11. These sensors detect pressure from the bending of piezo films. By increasing their sensitivity, they can detect the presence of an object in the immediate vicinity, because of static electricity. These technologies for humanlike appearance, behavior, and perception enable us to realize the humanlike androids. 9 Android Science 207 As discussed above, the framework of android science includes two approaches: the engineering approach and the scientific approach. The most typical experiment where the two approaches meet is the total Turing test" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001981_ijmpt.2016.079200-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001981_ijmpt.2016.079200-Figure1-1.png", "caption": "Figure 1 Cutting force acting on turning cutting tool", "texts": [ " The tool is held rigidly in a tool post and moved at a constant rate along the axis of the bar, cutting away a layer of metal to form a cylinder or a surface of more complex profile (Trent, 1991). In turning process, the workpiece is rotated and the cutting tool travelling to the left or right, removes a surface layer (chip) of the workpiece material. CNC Turning is a process that is used for precision machining to produce cylindrical components like hubs, rods, pulleys, etc., through metal cutting while using Computer Numerically Controlled lathes (Bhardwaj, 2009). Cutting forces are measured in three directions as shown in Figure 1. Three cutting forces components will be produced such as tangential force (Ft), which acts on the cutting speed direction, the feed force (Ff), which acts on the feed rate direction and the radial force (Fr), which acts on the direction which is normal to the cutting speed (Soukatzidis et al., 2005). It was observed that the cutting forces are directly depended on the cutting parameters such as cutting speed, feed rate, cutting depth, tool material, tool geometry, workpiece material and cooling types (Korkut et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure4-1.png", "caption": "Figure 4. Ventilated disc", "texts": [ " There are two types of disc: full discs and ventilated discs. The full discs, of simple geometry and thus of simple manufacture, are generally placed on the rear axle of the car. They are composed quite simply of a full crown connected to a \u201cbowl\u201d which is fixed to the hub of the car (Fig.3). The ventilated discs, of more complex geometry, appeared more tardily. They are most of the time on the nose gear. However, they are increasingly at the rear and front cars, upscale. composed of two crowns - called flasks - separated by fins (Fig. 4), they cook better than the full discs thanks to ventilation between the fins, which, moreover, promote convective heat transfer by increasing the exchange surface discs. The ventilated disc comprises more matter than the full disc; its capacity for calorific absorption is thus better. The literature review of the phenomena of braking shows that the principal request comes from the strong variations in temperature induced by the friction of the pads against the disc. Thermal load is defined as the temperature that causes the effect on braking, such as gradients of Temperature, and thermal cycles due to conduction and convection or radiation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001698_icelmach.2014.6960379-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001698_icelmach.2014.6960379-Figure3-1.png", "caption": "Fig. 3. 2D FEM model of PM generator.", "texts": [ " The most important parameters and design data of PM generator: \u2212 Nominal power: PN = 450 W, \u2212 Rated speed: nN = 750 rpm, \u2212 Nominal voltage: UN = 40 V, \u2212 Number of poles: 2p = 8, \u2212 Shaft height: H = 90 mm, \u2212 Winding: three-phase star-connected, distributed, \u2212 Position of magnets: surface mounted. Analysis of PMSM Vibrations Based on Back-EMF Measurements M. Bara\u0144ski, T. Jarek O 978-1-4799-4389-0/14/$31.00 \u00a92014 IEEE 1492 III. COMPUTER CALCULATIONS AND LABORATORY MEASUREMENTS Computational model of permanent magnet generator has been created in ANSYS Maxwell software. To minimize calculation time a two-dimensional model has been choosen. This simplification made it imposible to analyze vibration in Z-axis (parellel to the machine axis). Developed model of generator is shown in Fig. 3. Calculations and laboratory measurements have been carried out only at no-load condition \u2013 no load connected to the terminals of the machine. Vibrations of generator have been simulated as: \u2212 harmonic motion of the rotor relative to the stator in direction perpendicular to the axis of shaft (Fig. 4a), \u2212 oscilatory motion of the rotor relative to the stator around the axis of shaft (Fig. 4b). Case shown in Fig. 4a shows a possible state of bearing damage. Due to the fact that the simulated movements of the rotor are small (especially in the case of Figure 4a), an important aspect of the creation of the FEM model is to create a respectively dense finite element mesh particularly in the air gap" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000514_j.mechmachtheory.2019.103667-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000514_j.mechmachtheory.2019.103667-Figure2-1.png", "caption": "Fig. 2. Finite element model of a face-gear drive with supporting shafts.", "texts": [ " Points P1 and P2 are located at the intersection between the pinion axis of rotation and its corresponding front and back planes, respectively, whereas points W1 and W2 represent counterparts of points P1 and P2 for the face gear. The shafts of pinion and face-gear are designed as hollowed with outer and inner diameters represented by do1 and di1 for the pinion and by do2 and di2 for the face-gear (see Fig. 1). Points B1, B2, B3 and B4 are located at the position of the supporting bearings for the shafts. Both gears are mounted in overhanging configuration. The transmitted nominal torque of 2500 N m is applied at B2 on the pinion shaft, and the face-gear shaft rotation is restricted at B4. Fig. 2 shows all the components of the finite element model of the face-gear drive for determination of errors of alignment due to shaft deflections. Here, five pairs of contacting teeth have been used to avoid influence of the boundary conditions on the results and to take into account the load sharing between pinion and face-gear tooth surfaces. Facegear active tooth surfaces, which have larger curvature radii than the pinion tooth surfaces, have been defined as master surfaces, whereas the pinion active tooth surfaces have been defined as slave surfaces" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002475_s40430-020-02295-5-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002475_s40430-020-02295-5-Figure1-1.png", "caption": "Fig. 1 Interaction among the bearing components", "texts": [ " Here, the radiation noise model of a FCACBB is established based on the dynamic characteristics of bearings and acoustic theory. A number of indicators are used to characterize the sound field characteristics, and then, the calculation of the radiation noise is conducted. Assuming that the center of mass of each bearing component is consistent with the geometrical center, the movement of the cage is guided by the inner ring. To better analyze the vibration and noise characteristics of the high-speed FCACBB, as shown in Fig.\u00a01, the interactions among the bearing components are first analyzed. Then, the differential equations of vibration for each bearing component are established. The radiation sound pressure of a FCACBB is further calculated. In Fig.\u00a01, the inertial coordinate system {O; X, Y, Z} is fixed, and coordinate origin O is fixed to the initial center of the bearing. The X-axis represents the bearing rotation axis, which is parallel to the ground. The Y and Z axes represent the horizontal radial and vertical radial directions, respectively. The following notation is used to describe the components: ball (b), inner ring (i), outer ring (o), cage (c), cage pocket (p), and ordinal number (j) for balls. The vibration characteristics of each component are analyzed below [33, 34]. Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:311 1 3 Page 3 of 16 311 As an important component of an angular contact ball bearing, the balls contact all components, i.e., the cage, inner, and outer rings. It is supposed that all ceramic balls are equal in mass and size, and the sizes of the cage pockets, which are distributed uniformly along the circumference, are the same. In Fig.\u00a01, \u03b1ij and \u03b1oj are the contact angles between the jth ceramic ball and inner and outer raceways; Qij and Qoj are the normal contact forces between the jth ceramic ball and inner and outer raceways, respectively; T\u03b7ij, T\u03b7oj, T\u03beij, and T\u03beoj speed of the inner ring; \u03d5cj is the position angle of the jth ceramic ball relative to the cage; \u03c6ij is the position angle of the jth ceramic ball relative to the inner ring. The normal contact forces can be obtained through Hertz contact theory [35], and the traction forces can be calculated [36] by \u03b7VcSc/hc, where \u03b7 is the lubricant viscosity; Vc is the relative velocity of the ball and raceways; Sc is the contact area; and hc is the film thickness", "bzj are the traction forces of the contact surface between the jth ceramic ball and raceways; Qcj is the collision force between the jth ceramic ball and cage; Qbxj, Qbyj, and Qbzj are the decomposition components of Qcj, which are acting on the jth ceramic ball; Gbxj and Gbyj are the decomposition components of the gravity of the jth ceramic ball; P\u03b7j and P\u03bej are the friction forces acting on the surface of the jth ceramic ball, including rolling and sliding friction forces; Fbxj, Fbyj, and Fbzj are the components of the hydrodynamic force acting on the center of the jth ceramic ball; F\u03b7ij, F\u03b7oj, F\u03beij, and F\u03beoj are the hydrodynamic friction forces at the lubricant inlet of the contact zone of ball and raceways; Fwj is the aerodynamic resistance acting on the jth ceramic ball by gas\u2013oil mixture; Jxj, Jyj, and Jzj are the components of the moment of inertia for the jth ceramic ball; \u03c9xj, \u03c9yj, and \u03c9zj are the components of the spin angular velocity of the jth ceramic ball; ec is the relative eccentricity of the cage center; n is the rotation where DW is the diameter of the ceramic ball; mb is its mass; x\u0308bj , y\u0308bj , and z\u0308bj are the displacement accelerations of the barycenter of the jth ball; ?\u0307?xj , ?\u0307?yj , and ?\u0307?zj are the components of the spin angular acceleration of the jth ball; \u03c9bxj, \u03c9byj, and \u03c9bzj are the angular velocities of the jth ball; ?\u0307?bxj , ?\u0307?byj , and ?\u0307?bzj are the angular accelerations of the jth ball; ?\u0307?bj is the orbit speed of the jth ball; Ib is the moment of inertia of the ball. During the operation of the bearing, the cage contacts only the ball, and friction and impact are generated. In Fig.\u00a01, Fcx, and Fcy are the components of the hydrodynamic force, Fc, acting on the cage; Mcz is the friction moment acting on the cage. The differential equation for the vibration of the cage can be described as Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:311 1 3 311 Page 4 of 16 where mc is the mass of the cage; dm is the pitch diameter of the bearing; N is the number of the ceramic ball; Qcxj, Qcyj, and Qczj are the decomposition components of Qcj, which are acting on the cage; x\u0308c , y\u0308c , and z\u0308c are the displacement accelerations of the barycenter of the cage; \u03c9cx, \u03c9cy, and \u03c9cz are the angular velocities of the cage; " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002913_j.optlastec.2020.106727-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002913_j.optlastec.2020.106727-Figure3-1.png", "caption": "Fig. 3. Diagrammatic drawing of overlapped areas during rotating laser irradiation.", "texts": [ " For laser beam overlap, the time interval was approximately equal to the rotation period, which was in the order of milliseconds. Considering the extremely fast cooling rate during laser irradiation (approximately 103\u2013108 K/s) [24,25], the thermal effects of repetition irradiation were dominated by a laser-beam overlapping process. Therefore, the laser-ring overlapping effects were ignored in the temperature calculation. The combination of rotating laser beam and moving laser ring produced non-uniform overlapped areas, as shown in Fig. 3 (a). The concept of diagrammatic drawing here was employed to compare the overlapped areas at each position within one rotation period separately. In this case, the laser beam returned to its previous spot with horizontal displacement owing to the movement of the laser ring. When the laser beam reached the middle position of the laser ring, the overlapped area contained both the traveled and covered areas, denoted as A1\u2019 and A1 in Fig. 3 (a), respectively. Similarly, the overlapped area at top position was composed of A2\u2019 and A2. It is evident that the scanning laser beam traveled through a previously irradiated area within a rotation period, causing repetition irradiation. The same effects took place at any position of the laser ring because the scanning speed vb was much faster than the moving speed vr. Therefore, a rotating laser beam creates different overlapped areas during irradiation, even if the laser ring moves at a certain speed", " Note that the overlapped areas at each repetition irradiation decreased with the movement of the laser ring. In a study by Ashby and Easterling [16], the interaction time of the laser beam was defined as the S. Zhuang et al. Optics and Laser Technology 135 (2021) 106727 beam radius divided by the scanning speed owing to the low-power density around the edge of the laser beam. To decide the number of repetition irradiation n, the varying overlapped areas during repetition irradiation were assumed to possess the same thermal effects until the overlapped width exceeded the beam radius Rb, as shown in Fig. 3 (b). It is evident that the overlapped width at middle position started to exceed the beam radius when the laser ring displacement l1 equaled Rb. However, a longer displacement l2 was required for top and bottom positions. Here, the top/bottom and middle positions were selected to calculate the peak temperature distribution as representatives of maximum and minimum temperature values, respectively. For fixed rotation period tr and moving speed vr, the number of repetition irradiation n is expressed as follows: n = l vrtr (13) where l is the displacement of the laser ring when the overlapped width of the laser beam equals the beam radius (m); vr is the moving speed of the laser ring (m\u22c5s\u2212 1); and tr is the rotation period of the laser beam (s). According to Fig. 3 (b), the displacement of the laser ring l at middle position is given by the following expression: The top/bottom position has an l expressed as follows: Note that the interval time ti in rotating laser treatment is a fixed value and equals the rotation period tr. Therefore, Eq. (9) can be simplified as follows: Tn\u2212 1(z) = T0 +(n \u2212 1) { AP 2\u03c0\u03bbvb \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 [tr(tr + t0) ] \u221a exp [ \u2212 (z + z0) 2 4\u03b1tr ]} (14) Then, the peak temperature distribution can be obtained from Eqs. (11), (12), and (8)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000822_978-1-4939-0676-5_15-Figure15.8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000822_978-1-4939-0676-5_15-Figure15.8-1.png", "caption": "Fig. 15.8 Geometry of diffusion at a microband", "texts": [ " In the long-time limit, the current can be predicted by equation T2,11, which still contains the parameter \u03c4. Therefore, it is not a steady state. However, the current shows a logarithmic dependence of the time and declines rather slowly, so that it assumes a quasi-steady state. A band microelectrode is a two-dimensional diffusion system in which the length of the electrode is very much larger than the width. The coordinate system used to treat the diffusion problem at this geometry is shown in Fig. 15.8 and highlights that diffusion essentially occurs only along the x- and z-axes. This, as for the microdisks, causes the current density to be distributed nonuniformly and is especially infinite at the edge of the electrode. Analytical expressions for the current-time curves have been derived for long and short times.11 These are in the form of a series expansion, difficult to handle. A more convenient relationship that represents the two expansions as a closed form has also been derived (Table 15" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001392_elektro.2014.6848921-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001392_elektro.2014.6848921-Figure3-1.png", "caption": "Fig. 3. Detail of the oversaturated rotor tooth - broken rotor bar is on the right side of the over-saturated tooth", "texts": [ " Because of their small air-gap, attention has to be paid especially in case of induction machines. In this paper only SE is taken into account. B. Broken Rotor Bar Broken rotor bar occurs especially in machines with welded or soldered squirrel cages. Thermal expansion or torque ripple impact can cause the violation of the weld between rotor bar and ring. Current termination in broken rotor bar results in leakage flux extinction in the area of that bar. Then rotor tooth next to broken bar in direction of rotation is oversaturated (see Fig.3) and relatively large radial force originates in the area. This force rotates around the air-gap and disrupts the radial force balance in the machine. Consequently the stator core or frame vibrations can occur. There is also higher current in the bars neighboring with the broken bar, which leads to the temperature rise in these bars. III. FINITE ELEMENT MODELS DESCRIPTION Forces caused by modeled fault states are determined by 20 FE model of 4 pole, 1 1 kW SIEMENS I LA 7 163-4AA I 0 induction motor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002581_tte.2020.3008302-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002581_tte.2020.3008302-Figure2-1.png", "caption": "Fig. 2. Dismantled commercial off-the-shelf PMSM with fractional slot concentrating windings and flux concentrating magnet topology", "texts": [ " In [2], the development of a geared linear EMA technology demonstrator is described. In Fig. 1, the demonstrator actuates an abstracted aileron control surface via a hinge kinematics in a redundant active/active configuration [3]. The resolver is needed for velocity measurement and the (inverse) Park transformation converting current measurement and voltage reference values from a stationary three-phase into a rotating dq-reference frame and vice versa. A sensorless control mode would considerably reduce the length and hardware complexity of the motor in Fig. 2 by substituting the resolver with an estimation algorithm. Alternatively, such a control mode can provide an additional angle value, enabling fault tolerance through fast reconfiguration in case of a sensor failure. This mode can enhance redundancy concepts, such as that proposed in [4], providing novel health monitors and a new fail-operational mode. The properties of the PMSM are listed in Appx. A. In Sect. II, methods from literature for sensorless angle estimation are reviewed and the challenges in the critical low-speed region that is highly typical for flight control applications, are explained", " Downloaded on July 15,2020 at 14:30:50 UTC from IEEE Xplore. Restrictions apply. 2332-7782 (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. Alternatively, the utilization of an FEA tool for inductance identification is explained in Sect. III-B. In Sect. IV, both methods are applied independently to determine the anisotropy properties of the PMSM in Fig. 2. Beyond a well-known current dependency of the anisotropy coefficient, this paper systematically investigates a significant correlation with the rotor angle. In Sect. V, the test and simulation results are compared and the strengths and limitations of each method discussed. The implications on the feasibility of a sensorless control mode for the EMA demonstrator are highlighted and an outlook provided. The general model of a PMSM in a d/q-reference frame is given by ( ud uq ) = [ R \u2212\u03c9elLq \u03c9elLd R ]( id iq ) +[ Ldd Ldq Lqd Lqq ] \u02d9(id iq ) + ( 0 \u03c9el\u03a8PM ) (1) with the direct/quadrature axis voltage ud/q , current id/q , stator resistance R, electrical rotor velocity \u03c9el, absolute inductance Ld/q , differential inductance Ldd/qq, differential cross-coupling inductance Ldq = Lqd, and the permanent magnet flux \u03a8PM " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000868_gt2015-43971-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000868_gt2015-43971-Figure7-1.png", "caption": "Figure 7: IN-718 impellers with laser-consolidated blades built on premachined substrate (Left one with as-consolidated surface and the right one with sand blasted surface).", "texts": [ " An impeller shape (Figure 5) was selected to demonstrate the capability of LC process using a 5-axis CNC motion system. The impeller has a diameter of about 77 mm and height of about 26 mm. There are 9 long blades and 9 short blades uniformly distributed. A profiled substrate was pre-machined using a CNC lathe (Figure 6) and was mounted on a 5-axis CNC motion system using a simply designed fixture. Using a 5-axis CNC motion system to deal with large tiltrotation movement, LC was successfully conducted to build blades on the pre-machined substrate to form an integrated impeller. Figure 7 shows an IN-718 impeller (left) with its blades directly built up on a pre-machined substrate using LC process. After sand blasting, laser-consolidated blades and premachined substrate show consistent surface finish (Figure 7, right side). It is evident that this novel process produces high quality, fairly complex shapes directly from CAD models. The bond 4 Copyright \u00a9 2015 by Her Majesty the Queen in Right of Canada Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use between the LC features and the existing component is metallurgically sound, without crack and porosity [7, 10]. Compared to conventional welding process, the heat input from LC process to the substrate is minimal, resulting in a very small heat affected zone (several tens micrometers)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure19.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure19.2-1.png", "caption": "Figure 19.2 Low-pressure die-casting machine", "texts": [ " The dies are designed so that the moving half contains the male shape, and so, due to shrinkage of the metal, the casting clings to this half. When the moving platen nears the end of its stroke, striker pins on the ejector mechanism contact the underside of the fixed top platen, pushing the ejector pins which release the casting from the die. In some cases a stripper ring is used to release the casting, but the principle is the same. The moving platen is then reversed to close the two halves of the die ready for the next casting to be produced. Fig.\u00a019.2. The dies are cheaper than for high-pressure methods since they can usually be made of cast iron, but they are sometimes more expensive than for gravity die-casting. The capital outlay for machines is higher than for gravity but lower than for the high-pressure method. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 19 Die-casting and metal injection moulding 19 283 As the plunger is depressed, it covers up the filling port and so prevents molten metal escaping back into the crucible" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure2.13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure2.13-1.png", "caption": "Figure 2.13 Scrapers", "texts": [ " The purpose of scraping is therefore to remove high spots to make the surface flat or circular, and at the same time to create small pockets in which lubricant can be held between the two surfaces. Surface plates and surface tables are examples of scraping being used when flatness is of prime importance. Examples where both flatness and lubricating properties are required can be seen on the sliding surfaces of centre lathes and milling, shaping and grinding machines. The flat scraper, for use on flat surfaces, resembles a hand file thinned down at the point, but it does not have any teeth cut on it, Fig.\u00a02.13. The point is slightly curved, and the cutting edges are kept sharp by means of an oilstone. The scraper cuts on the forward stroke, the high spots being removed one at a time by short forward rocking strokes. The flatness is checked with reference to a surface plate. A light film of engineer\u2019s blue is smeared evenly on the surface plate, and the surface being scraped is placed on top and moved slightly from side to side. Any high spots show up as blue spots, and these are reduced by scraping. The surface is again checked, rescraped and the process is repeated until the desired flatness is obtained", "11 Correct angle of chisel well back at the end of the handle, not at the end nearest the head. Never allow a large \u2018mushroom \u2019 head to form on the head of a chisel, as a glancing blow from the hammer can dislodge a chip which could fly off and damage your face or hand. Always grind off any sign of a mushroom head as it develops, Fig. 2.12. The same procedure is used on internal curved surfaces, using a half-round scraper slightly hollow on the underside, to prevent digging in, and with a cutting edge on each side, Fig. 2.13. The reference surface in this case is the shaft which is to run in the curved surface and which is smeared with engineer\u2019s blue. Entry of the shaft in the bearing indicates the high spots, which are removed by scraping, and this process is repeated until the desired surface is produced. The three-square or triangular scraper, Fig. 2.13 is commonly used to remove the sharp edges from curved surfaces and holes. It is not suited to scraping internal curved surfaces, due to the steeper angle of the cutting edges tending to dig into the surface. However, the sharp point is useful where a curved surface is required up to a sharp corner. The engineer\u2019s hammer consists of a hardened and tempered steel head, varying in mass from D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 2 Hand processes 2 39 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure19-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure19-1.png", "caption": "Fig. 19. Application of eccentric axial loads.", "texts": [ " Three angles \u03b1, \u03b2, \u03b3 and a distance d are used to define the direction and position of the connecting members, where \u03b1 is the angle between the x axis and the projection of the member on the x-y plane, \u03b2 is the angle between the member direction and the x-y plane, \u03b3 is the rotation angle of the members along its own centre line, and d is the distance from the centre point of the node to the closer end section of the member. The width and the height of all connecting members are 80 mm and 180 mm respectively. The geometry of the node is shown in Fig. 19. It should be noted that the test rig design procedures discussed here are for testing nodes under individual loading case. Similar procedures can be applied to the design for combined loading case. The first step is to calculate the eccentricities and apply the axial loads at the eccentric centres of the connecting faces (Fig. 19). In this example, constant eccentricities of 30 mm are applied in both in-plane and out-of-plane directions. In the second step, eye-connection plates are designed to be vertically mounted on the connecting members as shown in Fig. 20. As discussed before, the positions of bolt holes on these plates should be determined with the consideration of the bolt looseness, so that after initial movement, the bolts should be placed in the right positions as they are designed for. As can be seen in Fig. 20, the vertical plates are parallel to the connected members\u2019 centre-lines" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure3.8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure3.8-1.png", "caption": "Fig. 3.8 Collision analysis of a gripper pose. The yellow volumes are unclassifiable threat volumes, as the upper and the lower surface of the gripper are below the scanned surface. Therefore, the areas can be occluded collision or occluded free space. The red areas are collision volumes, as the upper part of the gripper is above and the lower part below the scanned surface", "texts": [ "4 Pose hypotheses generation using a \u2018birthday attack\u2019like approach. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 Figure 3.5 Scan of piston rods lying on a table (SICK LMS400) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 Figure 3.6 Visualization of the occurring problem when sheet metal parts are scanned from a single point of view . . . . . . . 21 Figure 3.7 Rotational and translational invariant features of a tripole . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22 Figure 3.8 Threat volume versus collision volume . . . . . . . . . . . . . . . . 26 Figure 3.9 The gripper used for all experiments . . . . . . . . . . . . . . . . . 26 Figure 3.10 Key Grasp Frame definition . . . . . . . . . . . . . . . . . . . . . . . 27 Figure 3.11 Example for the optimal grasp pose estimation algorithm only using the palm of the gripper . . . . . . . . . . . . 28 Figure 3.12 Models of the Mian data set . . . . . . . . . . . . . . . . . . . . . . . 29 Figure 3.13 Examples of models of the data set of industrial parts", " Whenever a pixel of the upper face of an endeffector part is above the scan, but the same pixel of the lower face is located below, a collision in that pixel of the image is present. With the value of the differences of the pixels and the size of the pixels a collision volume can be calculated for this pixel. Whenever both surfaces are located below the scan, no clear evaluation can be done for that pixel because the gripper could be in occluded free space or in occluded collisions. The resulting unclassifiable volume is called threat volume (see Fig. 3.8). Collision and threat volume result in an overall collision volume Vc = \u2211 Part \u2211 i sx sydiwi . (3.12) Here, sx and sy are the pixel dimensions, di is the height difference of gripper surface and scan surface in the colliding pixel i and wi is a weight which can be chosen to distinguish between collision and threat volume. Furthermore,wi can give different weights to different parts of the gripper (Fig. 3.9). At this point, it was shown how single grasp poses can be efficiently evaluated. To further enhance the performance, the approach is extended to work on predefined grasp regions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001498_j.wear.2015.11.008-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001498_j.wear.2015.11.008-Figure3-1.png", "caption": "Fig. 3. Dimpled thrust bearing with phyllotactic pattern (a) prototype, (b) geometrical description.", "texts": [ " The lubricating dimples were positioned according to the phyllotactic pattern on the thrust bearing to form a textured working surface. In theory, these dimples were placed in a manner that conserves lubricating oil while improving the friction properties of the thrust bearing due to environmental adaptation. To validate this hypothesis, we tested a set of phyllotactic dimpled thrust bearings. Dimple coordinates were calculated using the phyllotaxis equations provided above at angle \u03b8 of 137.508\u00b0. A prototype of the dimpled thrust bearing, with its dimensions labeled, is shown in Fig. 3. While computing the dimple coordinates, differing phyllotactic coefficient c and dimple diameter d were selected (from 2 to 6 and from 2 to 4 mm, respectively.) These c and d values produced dimples with proportion of 10\u201320% of the area of the entire flat surface. The empirical value of dimple depth \u03b4, 0.03\u20130.08, was selected. (a) photograph, (b) phyllotactic pattern. Thrust bearings were fabricated with CuZn25Al6Mn4 material. The specimens had identical dimensions: internal diameter d0 of 13.2 mm, external diameter D of 40 mm, and thickness H of 3 mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002511_s0036029520060099-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002511_s0036029520060099-Figure2-1.png", "caption": "Fig. 2. Schematic diagram", "texts": [ " For the implementation of selective laser melting and sintering technologies, OOO Moscow Laser Technology Center developed and manufactured an SLP-110 unit. The appearance of the installation and its design are shown in Fig. 1. The installation can be used for fully automatic single and small-batch production of complex products from special metal powders, namely, stainless and tool steels, aluminum alloys, nickel alloys, titanium alloys, and cobalt\u2013chromium compositions. The manufactured part size is 110 \u00d7 110 \u00d7 110 mm. The operation scheme of the installation is shown in Fig. 2. A powder is fed to the bench using a hopper with a platform moving along the vertical. The powder applied onto the bench is leveled off with a roller. The powder is melted by a laser beam moving along the 649 surface with a mirror scanner. The minimum single layer thickness is 20 \u03bcm, and the rate of building of parts is up to 10 cm3/h. One of the important conditions is the creation of a protective environment during the growth process. For this purpose, the installation has a sealed chamber with a controlled atmosphere" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000566_ab5b65-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000566_ab5b65-Figure8-1.png", "caption": "Figure 8. (a) Experimental setup, (b) schematic structure of the planetary gearbox.", "texts": [ " In figure\u00a07(b), even that most of the noise is removed by MTSA, the useful features are still not evident in figures\u00a07(b)-2 and (b)-3. And yet for that, by virtue of PRPCA, the fault anomaly is faultlessly separated under all the three levels of noise as shown in figure\u00a0 7(c), which validates its excellent performance for periodic anomaly extraction. In order to present the validity of the proposed method in practical applications, different types of tooth defect on a planetary gearbox are introduced for the analysis in this experiment. The experimental setup is shown in figure\u00a08(a), which consists of a servo motor, a magnetic break for loading and a planetary gearbox with an internal structure as given in figure\u00a08(b). Besides, two Heidenhain rotary encoders with 5000 pulses per revolution are equipped on this experiment rig to measure the angular displacements of input shaft and output shaft respectively. In this paper, the output encoder signals were collected for analysis by an IK220 counter card at a sampling rate of 5000 Hz. The computational formulas of characteristic frequencies (CFs) are listed in table\u00a02, where Np = 3 is the number of planet gears equally spaced in the gearbox, f i represents input frequency and Zp , Zs, Zr denote the number of teeth on planet gear, sun gear and ring gear, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001851_s12206-016-0614-7-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001851_s12206-016-0614-7-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of the taper leaf spring of tandem suspension.", "texts": [ " The present paper proposes a way to calculate dynamic spring rate for leaf spring and provides the details about that formula. The simulations are conducted with the same condition as the test. The hysteretic characteristics and relative error of dynamic spring rate from the test are compared with the ones from the simulation for the validation of the high precision of the modified model. This study develops a high-precision model of the taper leaf spring used in the tandem suspension of a multi-axle vehicle. Fig. 1 shows a schematic diagram of the taper leaf spring of tandem suspension when the taper leaf spring is installed. The intermediate and the rear axles are connected by the taper leaf spring, and both ends of the taper leaf spring are freely supported at the intermediate and rear axles. The middle of the taper leaf spring is connected to a frame through a U-bolt. The taper leaf spring of tandem suspension does not have a shackle, joints, and bushing elements compared with the traditional taper leaf spring" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001715_000393788-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001715_000393788-Figure2-1.png", "caption": "Fig. 2. Test set-up for resistance measurement. 1 = horizontal steel shaft ; 2 = tube; 3 = vertical telescopic rod; 4 = force transducer ; 5 = horizontal connecting rod with spring system.", "texts": [ " swimmer and with a sound-meter in order to locate the contraction in the arm movement as shown in figure 1. The resistance tests were carried out in a towing tank (200 x 4 x 4 m) of the Netherlands Ship Model Basin. This technique has been applied to swimmers by \u039f\u03bfs\u03a4\u03b5\u03b1\u03bd\u03b5\u03c4.\u03bf and RuKEN [7], V\u0391\u039d MANEN and RuKEN [8] and by DE GoEDE et al. [3]. The basin walls are fitted with rails on which a towing carriage can be driven electrically at different, accurately adjustable speeds. The speed of the cariage is measured by a photo-electric cell system, and recorded on direct developing photo-strip chart. Figure 2 shows the test set-up for the measurement of forces acting on the moving body. The kinematographical analyses show important differences in the body position. The polo stroke is characterized by a lordosis in the lumbar region and an elevation of the shoulders. The light-trace analyses show a longer gl\u00edd- D ow nl oa de d by : U ni ve rs it\u00e9 R en \u00e9 D es ca rt es P ar is 5 19 3. 51 .8 5. 19 7 - 4/ 28 /2 01 8 10 :2 9: 33 P M ing period and consequently a longer arm cycle in the competition crawl. Biomechanically speaking, the arm movement patterns are very similar in both strokes" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000943_1.4030630-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000943_1.4030630-Figure16-1.png", "caption": "Fig. 16 Schematic diagram of the receiving area of the transducer", "texts": [ " fr (10) where c is the speed of the bearing cage and fr is the repetition frequency of the ultrasound pulse. Based on the relative motion between the roller and the transducer, the displacement (dr) can be regarded as the focal spot displacement. Under low-speed and high repetition frequency, the focal spot displacement can be much smaller than the focused beam diameter. This will bring about huge errors in measuring the results by the overlapping regions contained by any two consecutive measuring points, as shown in Fig. 16. So, the high PRR is not directly applicable for the low shaft speed. Therefore, for a successful measurement, the focus size should be less than the size of the contact bearing measured, and the focus size should also be less than the focal spot displacement. A new ultrasonic pulser-receiver was described and used for the measurement of lubricant-film thickness in a cylindrical roller bearing (type N220). This device provided a maximum pulse repetition frequency of 100 kHz, so that the ultrasonic measurements could be used for high shaft speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000588_9783527822201.ch15-Figure15.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000588_9783527822201.ch15-Figure15.4-1.png", "caption": "Figure 15.4 Schematic representation of the growing mechanism. The sequence (a\u2013c) shows successive addition of new layers of material at the apical level. Deposition of a new layer creates the force required for the penetration of the tip into the soil.", "texts": [ " The result is a lower friction path that helps in soil penetration reducing the total energy needs [11]. The growing mechanism in [12] can be considered as a miniature burrowing system that creates the tunnel structure while it moves forward by growing. This system has a cylindrical body that consists of a rotating deposition head that moves new material from the top to the growing zone and shapes a tubular structure out of its body. The body, fixed with the soil, supports the further construction of the body and the axial movement of the deposition head into the soil (Figure 15.4). The rotary motion of the deposition head, converted into a linear motion at the tip, provides the force for penetrating soil. The force generated by this screw-like mechanism produces an ideal mechanical advantage: MAideal = \u03c0D d = tan\u22121 \ud835\udefc (15.1) Assuming that the diameter of the filament (d) remains constant during the growth process (i.e. one deposition cycle of the tip penetrates by a distance equal to the filament diameter), D is the mean diameter of the tubular structure and \ud835\udefc the lead angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000709_psc49016.2019.9081455-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000709_psc49016.2019.9081455-Figure10-1.png", "caption": "Fig. 10. Field windings replaced with PMs [9].", "texts": [ " This categorization is illustrated in Fig. 8. In this section, different types of Maglev motors are explained. A. Linear Synchronous motors Linear synchronous machines have either a passive or an active track. In motors passive tracks, as shown in Fig. 9, both of the field and armature windings are placed in the train and by changing the reluctance of the track, polarity of the motor\u2019s flux is changed and the train is moved [8]. In a similar topology, field windings are replaced with PMs, as shown in Fig. 10. The PMs are placed between the armature windings and in the magnetic loop [9]. In order to improve the performance of the motor under fault conditions, one must reduce the mutual inductance of phase windings with each other. To that end, the design of Fig. 11 is proposed. In this design, when the current of one phase increases due to a fault, the side phases are minimally affected [10]. Another method of changing the reluctance of the track is using the zigzag configuration as shown in Fig. 12 [11]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001170_0278364914551773-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001170_0278364914551773-Figure2-1.png", "caption": "Fig. 2. Wheel slip. The true velocity of a wheel is the sum of an idealized (non-slip) velocity and a slip component.", "texts": [], "surrounding_texts": [ "Consider Figure 3. For illustration purposes, we will allow a single degree of articulation between the vehicle and contact point frames. Let us define the following frames of reference: w, world, fixed to the environment; v, vehicle (body), fixed to some point on the vehicle whose motion is of interest; s, steering, positioned at the hip/steering joint, moves with the boom (if any) to the wheel; c, contact point (wheel), moves with the contact point, has the same orientation as the steering frame. Each frame has a default associated Cartesian coordinate system, based on the axes shown in the figure. The axes in all such figures are x and y where x must be rotated 90 to be coincident with y. In this example, we will assume rolling terrain of surface curvature that is less than the wheel surface curvature, so the wheel contact point is fixed with at UNIV OF PITTSBURGH on February 10, 2015ijr.sagepub.comDownloaded from respect to the steering frame. We will assume conventional wheels. 3.3.1. Inverse kinematics. The solution to the inverse kinematics problem is a special case of Equation (53). In our case, we have chosen the case where v *v s = 0 but v *v s 6\u00bc 0. On rolling terrain where the bottom of the wheel remains in contact, we also have v *s c = 0, and this equation reduces to vw c = vw v rv c \u00d7 vw v rs c \u00d7 vv s \u00f055\u00de We will call this the offset wheel equation in matrix form. To use this, we need to know the rate of steering rotation vv s with respect to the vehicle. This is usually measurable or otherwise known to be zero if there is no steering rotation. The other vectors r *v c and r *s c are known vehicle dimensions. Also, in the case where the steering axis is coincident with the contact point, we can align frames c and s and then the last term will vanish. 3.3.2. Wheel steering control. Controlling such a wheel involves steering it and driving it around its axle. For steering, doing so in a kinematically correct fashion based on Equation (55) seems surprisingly difficult. The wheel steering angle g must be chosen such that the x axis of the c frame is aligned with the wheel velocity vector. Equivalently, the wheel velocity in the y direction of the contact point frame must vanish. Once this is so, the steering rate vv s is predetermined to be the time derivative of the necessary steering angle. The component of r *v c expressing the steering offset r *s c must also be consistent and all three of these quantities must be consistent with vw c which is also unknown because it depends on the unknown steering angle. In other words, everything other than gross motion and the position of the steering frame, is unknown. The solution to this dilemma is to recognize (a) that we need only the direction of the wheel velocity to steering correctly, (b) that the direction (not the magnitude) of the c frame velocity with respect to the world must always be parallel to that of the s frame if the link joining them is rigid, and (c) that the velocity of s the frame is completely determined by the velocity of the vehicle frame. That is, we can solve the problem recursively by first propagating the known quantities vw v and vw v as far as possible up the kinematic chain to the wheel. The steering frame velocity is given by vw s = vw v rv s \u00d7 vw v \u00f056\u00de When this velocity is expressed in vehicle coordinates, the angle of the contact point velocity with respect to the vehicle body must be parallel to it in any coordinates: g = arctan 2\u00bd(vvw s )y, ( vvw s )x \u00f057\u00de The wheel can be steered to this angle, and it will then meet the conditions for rolling without lateral slipping. In the more general case of motion over rough terrain, the contact point frame must move over the wheel and, for convenience in requiring no terrain penetration, must remain oriented with the terrain normal. Such a moving contact point frame violates two of the assumptions above. In that case, one can define a wheel axle frame a oriented parallel to the steering frame, with respect to which the contact point moves. Then, the contact point velocity va c relative to the wheel can be subtracted from vw c to produce vw a and the above technique can be used. 3.3.3. Wheel speed control. Once the steering angle is known, the steering rate becomes known from differentiation, and the geometry of the whole kinematic chain becomes known. We could compute the component of vw c oriented along the x axis of the frame but we often already know the y component is zero, so we can simply compute the magnitude of the velocity. Furthermore, we can compute it in any coordinates. In the vehicle frame, the contact point velocity is computed as vvw c = vvw v rv c \u00d7 v vw v Rv srs c \u00d7 v vv s \u00f058\u00de The wheel translational speed is then vw c = ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi vvw c 2 x + vvw c 2 y q \u00f059\u00de Assuming no longitudinal slip, this can be converted to angular velocities of the wheel (about its axle) using vwheel = vw c =rwheel \u00f060\u00de where rwheel is the wheel radius. Extensions to include known longitudinal slip are straightforward. The sign of the wheel angular velocity may need to be adjusted depending on the conventions in effect for the positive sense of rotation. 3.3.4. Wheel velocity sensing. Suppose we have measurements of wheel angular velocities vk and steering angles gk and we want to determine the linear and angular velocity of the vehicle. Typically, the wheel angular velocity and steering angles are sensed more or less directly. Under this assumption, the linear velocities of the wheels are derived by inverting Equation (60): vk = rkvk \u00f061\u00de Assuming no slip, the wheel velocity components are directed along the x axis of the contact point frame. In the body frame, they are (vk)x = vk cos (gk) (vk)y = vk sin (gk) \u00f062\u00de 3.3.5. Forward kinematics. On flat terrain, there are three unknowns in the body velocity (vw v ,v w v ) and measurements at UNIV OF PITTSBURGH on February 10, 2015ijr.sagepub.comDownloaded from provide only two constraints (gk,vk) per wheel. Therefore, we need at least two sets of wheel measurements in general to determine the motion of the vehicle frame. However, the problem then becomes overdetemined with at least four conditions on three unknowns. Steering axes cannot be perfectly aligned in practice, and wheel encoders will often not be consistent, so a best fit solution, which tolerates the potential inconsistency in the measurements, is needed in general. Consider again Equation (58), the wheel equation in vehicle coordinates: vvw c = vvw v rv c \u00d7 v vw v Rv srs c \u00d7 v vv s \u00f063\u00de This is of the form vvw c = Hv c (g) vvw v vvw v + Qs c(g) vvv s \u00f064\u00de Note that rv c depends on the steering angle g as does Rv s . If we knew the vehicle frame velocities, we could write Equation (63) for each wheel in order to compute the velocity of each wheel contact point in one matrix equation. Stacking all equations together and grouping the first two terms together produces a matrix equation of the form vvw c = Hv c (g) _x w v + Qs c(g) _g \u00f065\u00de Because the left-hand side and the steering angles are known, this is a set of simultaneous equations constraining the vehicle linear and angular velocity _xw v , and they can be solved using any suitable method, including the pseudoinverse. _xw v = Hv c (g) >Hv c (g) h i 1 Hv c (g) > vvw c Qs c(g) _g h i \u00f066\u00de For non-offset wheels Hv c simplifies, and Qs c disappears. This solution produces the linear and angular velocities that are most consistent (in a least-squares sense) with the measurements, even if the measurements do not agree." ] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure15.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure15.6-1.png", "caption": "Figure 15.6 Bending a splash guard for a centre lathe", "texts": [ " Simple bending is carried out by locally heating along the bend line, from both sides, until the material is pliable, using a strip heater. A strip heater can easily be constructed using a heating element inside a box structure, with the top made from a heat-resisting material. The top has a 5 mm wide slot along its centre, through which the heat passes, Fig.\u00a015.5. When the material is pliable, it can be located in a former and bent to the required angle, e.g. in making a splash guard for a lathe, Fig.\u00a015.6. The material can be removed from the former when the temperature drops to about 60 \u00b0C. Formers can be simply made from any convenient material such as wood. For volume production large-scale industrial machines are available which can have multiple bending areas and advanced heat control. Shapes other than simple bends can be carried out by heating the complete piece of material in an oven. To avoid marking the surface, the material can be placed on a piece of brown paper. The time in the oven depends on the type of material and its thickness, and time must be allowed for the material to reach an even temperature throughout" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001740_1077546315619071-Figure27-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001740_1077546315619071-Figure27-1.png", "caption": "Figure 27. Experimental set-up: (a) Test-rig; (b) Test bearings with a defect size of 500mm on the outer race; (c) Test bearings with a defect size of 500mm on the inner race (Shakya et al., 2013; Shakya et al., 2015).", "texts": [ " The validation of the methodology is also carried out on the vibration data available on an online database (CWRU). In this section, details of the experimental set-up for seeded defect are at City University Library on December 19, 2015jvc.sagepub.comDownloaded from discussed followed by discussion on results of the application of the proposed methodology on experimental vibration signals. Test setup for bearings with seeded defect The test set-up used for the seeded defect experiments is shown in Figure 27. Position of the test bearings, accelerometers, and provisions for the application of the radial and axial load are highlighted in the figure. Double row angular contact ball bearings (NBC make, AU1103M) are utilized for experimentation. The radial and axial loads are 25% and 3.75 % of the dynamic load capacity of the bearing. The speed of the shaft is kept constant at 1400 RPM. Line defects, in the form of a fine slit, are etched on the outer and inner race with the help of electro discharge machining. Six cases of bearing defect with three different defect sizes of 500 mm, 1000 mm, and 1500mm on outer and inner races are studied. The test bearings with the outer and inner race defects of 500 mm size are shown in Figure 27b and c respectively. Before data acquisition, an initial running-in procedure is carried out for all the tests. Twenty-five data sets were acquired for a particular size of defect. An Leuven Measurement Systems (LMS) make SCADAS MOBILE SCM01 data acquisition system with inbuilt anti-aliasing filters is used for vibration data acquisition with a sampling rate of 51.2 kHz. Vibration data is acquired using a B&K make 4508-001 accelerometer, which has a sensitivity of 10mV/g. The accelerometer is a piezoelectric vibration sensor that is separate from the LMS make data acquisition system", " The sensitivity analysis on effect of change of load shows that the proposed methodology of defect classification is largely insensitive to practical variations in operating conditions. On the other hand, the defect severity assessment (DSV values, column 6) is affected by about 6\u201320% with almost 100% increase in the load, depending on the type of defect (inner/outer race). Thus, within the reasonable changes in the applied load, the defect severity value may not change appreciably, even for the outer race defect. The sensitivity analysis could be done only on CWRU data, since the vibration data was acquired from the test rig (Figure 27) at constant load and speed. A methodology is proposed to differentiate between the outer and the inner race defect in bearings. The methodology is based on vibration data and subsequent processing in the time-frequency domain. Instantaneous energy density (IE) corresponding to the interaction of rolling element and defect is calculated through HHT. The pattern of the variation of IE differs in case of outer and inner race defects due to modulation process at every defect interaction with rolling elements" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002523_j.matpr.2020.05.322-Figure25-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002523_j.matpr.2020.05.322-Figure25-1.png", "caption": "Fig. 25. Carbon/Epoxy Resin Anti roll Bar \u2013 Strain plot.", "texts": [], "surrounding_texts": [ "In this research article reveals that the high amount of stress resistant on tubular bar when compared to the steel bar. is The ng and finite element analysis of anti-roll bar using ANSYS software, Mate- Fig. 23. Carbon/Epoxy Resin Anti roll Bar \u2013 Displacement Results. Fig. 24. Carbon/Epoxy Resin Anti roll Bar \u2013 Factor of Safety plot. counter move bar was displayed in CATIA and anticipated outcomes were gotten in ANSYS. The suspension planners have the adaptability to tune the taking care of properties to their demanding models. These recreations likewise gave qualities to key structure factors. This is practiced by choosing the correct mix of distance across and divider thickness. They can build the solidness of the bar, without the weight punishments ordinarily connected Please cite this article as: V. Mohanavel, R. Iyankumar, M. Sundar et al., Modelli rials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.05.322 with a stiffer strong Anti-move bar. The following handling performance directly affects the vehicle roll. Adding roll stiffness to the front tends to make the vehicle more under steering and increasing roll stiffness in general tends to increase peak attainable side forces. Hollow tubular bar is protect the maximum amount of stress compared solid bar. ng and finite element analysis of anti-roll bar using ANSYS software, Mate- Fig. 27. Carbon/Epoxy Resin Anti roll Bar \u2013 stress plot in Y \u2013 Axis. Table 2 Properties of Anti roll bar. Properties Solid Bar Tubular Bar Mass, grams 911.45 328.15 Volume, mm3 455726.86 164075.87 Surface area, mm2 92732.89 166013.31 Centre of mass, Lxx, (gmm2) 5966209.98 2157061.67 Centre of mass, Lyy 127883238.68 46052599.14 Centre of mass, Lzz 133803869.90 48182739.26 Moment of Inertia, Ixx 49905934.16 17980034.85 Moment of Inertia, Iyy 127883238.80 46052600.66 Moment of Inertia, Izz 177743594.20 64005713.96 Table 3 FEA results obtained from ANSYS software. Properties Solid Tube Outer Diameter, mm 24 24 Inner Diameter, mm 0 18 Polar Moment of Inertia, mm4 32,572 22,266 Torque, Nm 100 100 Stress at OD, MPa 64 108 Stress at ID, MPa 0 86 Shear Stress, MPa 171.88 107.43 Axial Stress, MPa 339.63 212272.9 Max Stress obtained from the results, N/mm2 89937.78 622267.7 Please cite this article as: V. Mohanavel, R. Iyankumar, M. Sundar et al., Modelling and finite element analysis of anti-roll bar using ANSYS software, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.05.322" ] }, { "image_filename": "designv11_22_0002077_fuzz-ieee.2016.7737765-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002077_fuzz-ieee.2016.7737765-Figure3-1.png", "caption": "Fig. 3: Omnidirectional Mobile Robot Schematics", "texts": [ " They argued that omnidirectional mobile robot fails to reach the goal because of the wheel slippage and tried to find parameters which affect the wheel\u2019s rotational direction with various surface conditions real experiments and optical flow information. In this paper, we constrained the environment to single surface. Because of technological improvements in simulation tools, repeated simulations with different parameters for checking odometry errors can be done in short time with less efforts. Omnidirectional mobile robot whose schematic shown in Fig. 3 was implemented in Webots. Using Webots environment instead of using real experiments, each unit movement of the robot is optimized using QEA. In the omnidirectional mobile robot, each of three wheels are located perpendicular to the robot\u2019s center of gravity, where i-th wheel is located \u03c6i degrees away from the front of the robot. From the dynamics of the robot, translational and angular velocities of the robot are related as follows:\u23a1 \u23a3\u03d51 \u03d52 \u03d53 \u23a4 \u23a6 = 1 r \u23a1 \u23a3\u2212sin(\u03b8 + \u03c61) cos(\u03b8 + \u03c61) L \u2212sin(\u03b8 + \u03c62) cos(\u03b8 + \u03c62) L \u2212sin(\u03b8 + \u03c63) cos(\u03b8 + \u03c63) L \u23a4 \u23a6 \u23a1 \u23a3x\u0307y\u0307 \u03b8\u0307 \u23a4 \u23a6 (5) or in short: [\u03d5i] = 1 r [W ][S\u0307i]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001436_1350650113519611-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001436_1350650113519611-Figure10-1.png", "caption": "Figure 10. Comparison of foil bearing surface between experimental and numerical results.", "texts": [ " Figure 9 shows the theoretical top foil shapes at the radius of 5.25mm for the initial condition and at w\u00bc 4.3N and 350,000 r/min, and the pressure distribution for the four pads. The rotor surface shifts downward by increasing the applied load, the top foil of each pad is deformed according to the rotor surface. The pressure distribution of each pad is somewhat different because of the difference of the initial shape of the top foil. It can be said that aerodynamic pressure was properly generated in the bearing clearance. Figure 10(a) shows the photograph after the experiment at w0\u00bc 4.9N and n\u00bc 350,000 r/min. Contact marks were observed at several positions, in particular at the outer trailing edge of Pad 1, 3, and 4. In Figure 10(b), 2D images of the bearing clearance for the four pads are shown. By comparing this figure to Figure 10(a), the numerical bearing clearance became less than 5 mm at positions where contact marks were observed. Figure 11 shows the numerical and experimental results of the shaft displacement for the test bearing when the rotor speed was varied. In this figure, the rotor displacement was set to be zero at the maximum applied load. In calculations, equations (4) and (5) for structural stiffness of bump foils were employed and the numerical results obtained showed good agreement with experimental data" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002673_j.matpr.2020.07.229-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002673_j.matpr.2020.07.229-Figure9-1.png", "caption": "Fig. 9. Contour plot of 1st Mode shape (Bending mode) of cantilever beam.", "texts": [], "surrounding_texts": [ "Harmonic analysis for each material was conducted for the first three mode shapes. To conduct the harmonic analysis a load of Please cite this article as: P. Agrawal, P. Dhatrak and P. Choudhary, Comparativ method, Materials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.07 35303.94 N was applied on the tip of the beam. After considering the aircraft to be of 2000 kg the above load was obtained [10]. The analysis was conducted in steps of 100 and the frequency response graph with respect to amplitude was obtained. e study on vibration characteristics of aircraft wings using finite element .229 Fig. 10. Contour plot of 2nd Mode shape (Lateral mode) of cantilever beam. Fig. 11. Contour plot of 3rd Mode shape (Twisting mode) of cantilever beam." ] }, { "image_filename": "designv11_22_0001601_s00170-016-8487-6-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001601_s00170-016-8487-6-Figure3-1.png", "caption": "Fig. 3 The geometrical parameters of ECAE die along with the deformation of initially square element abcd after passing through the die and changing into deformed element a b c d ]15 ]", "texts": [ " D0 2 \u00bc 2ln D1 . D0 \u00f01\u00de When the billet is extruded from the globe-shaped hollow into the exit channel, the same amount of effective strain is accumulated in the material. Therefore, total effective strain induced over the expansion and the extrusion stages can be expressed as below: \u03b51 \u00bc 2ln D1 . D0 \u22122ln D0 . D1 \u00bc 4ln D1 . D0 \u00f02\u00de Effective strain caused by the shear deformation is [15]: \u03b52 \u00bc 2cot \u03d5 . 2\u00fe \u03c8 . 2 \u00fe \u03c8cosec \u03d5 . 2\u00fe \u03c8 . 2 h i. ffiffiffi 3 p \u00f03\u00de which is the same plastic strain as that for ECAE (Fig. 3), where \u03d5 is the angle between the channels and \u03c8 is the outer corner angle of ECAE die. Considering the flow pattern of material in an Exp-ECAE process, the\u03c8 approximately equals to 90\u00b0. Finally, by superimposing two recent components, the total effective strain attainable by means of an Exp-ECAE can be formulated as: D0\u03b5Total \u00bc \u03b51 \u00fe \u03b52 \u00bc 4ln D1 . D0 \u00fe 2cot \u03d5 . 2\u00fe \u03c8 . 2 \u00fe \u03c8cosec \u03d5 . 2\u00fe \u03c8 . 2 h i. ffiffiffi 3 p \u00f04\u00de Experimental tests were performed using the AA6063 aluminum alloy [14]. Rods with 20 mm in diameter were purchased then machined into the billets with a diameter and length of 15 and 120 mm, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000709_psc49016.2019.9081455-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000709_psc49016.2019.9081455-Figure16-1.png", "caption": "Fig. 16. Train with superconductive linear synchronous motors [13].", "texts": [ " In this design, two pieces of ferromagnetic matterial, which are named auxiliary poles, are placed at the two ends of the train to reduce the leakage flux. To reduce the cogging torque of LSMs, design of Fig. 15 is proposed. In this design the method Phase Set Shift is implemented, in which neighbouring phase sets are shifted by a certain degree [12]. One of the main types of long armature motors is superconductive linear synchronous motors. In these trains, in order to reduce the consumed power by the on board excitation circuit, the field windings are made of superconductive materials. Fig. 16 demonstrates this method [13]. Recently, as high temperature superconductors are becoming more and more commercialized, the superconductive materials like YBCO and BSCCO are being tested as the material for the windings. This will simplify the cooling system. Fig. 17 compares the cooling system of a low temperature superconductive coil and a high temperature one. B. Linear Induction motors The type of linear induction motor which is used for Maglevs is short armature. As demonstrated in Fig. 18, LIMs are composed of on board armature windings and a conductive layer installed on the track and as the electrical energy needs to be transferred into the train, the electric connectors are necessary, limiting the speed of train" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002550_s12555-018-0849-4-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002550_s12555-018-0849-4-Figure1-1.png", "caption": "Fig. 1. The Photograph of Quanser table-mount 3-DOF helicopter.", "texts": [ " Finally, Section 5 concludes this paper. Some basic preliminaries will be introduced before presenting the controller design process. Firstly, in Section 2.1, the dynamics of the 3-DOF helicopter in the presence of actuator faults, system uncertainties and external disturbances are presented. Then, we will design a second-order differentiator in Section 2.2, while Section 2.3 describes fixed-time SMO. 2.1. The dynamics of 3-DOF helicopter The research object 3-DOF helicopter, which is shown in Fig. 1, has three degree of freedom, elevation, pitch and travel [33]. According to the structure of the 3-DOF helicopter, we know that the two propellers make the helicopter move around the fulcrum in the vertical direction, that is, the pitching movement. When the sum of the two lifts is greater than the sum of the gravity G, helicopter to pivot upward movement. The adjustment of the distance between the balance block and the fulcrum can change the gravity of the helicopter. Since the 3-DOF helicopter has two inputs and three outputs and only two of three axes can be controlled arbitrarily, elevation and pitch axes are investigated and the travel axis is set to free to exam the control scheme proposed in this work", " The pitch angle and elevation angle are measured by position encoders and then are sent to computer. Their derivatives are obtained by passing through a second-order low-pass filter module in the 3-DOF test-bed. According to the measurements and proposed FTC method, control inputs can be generated. The desired trajectories for elevation and pitch angles are chose as x1d = \u22120.1+ 0.1sin(0.08t) and x3d = 0.1sin(0.06t). System uncertainties are mainly from modeling errors. To test the performance of the proposed strategy, external disturbances are generated by using a fan (see Fig. 1) and it is located such that it can act during the time interval. 85% control input is implied to front motor to simulate the efficiency of front motor is degraded to 85% and it appears at t \u2265 30s. The parameters of proposed FTC scheme are given as \u03bb1 = 8, \u03bb2 = 5, \u03bb3 = 9.5, KE1 = [ 4 3.5 ], KE2 = [ 3.5 3.5 ]. It can be seen from Figs. 2-5. that pitch angle is more sensitive than elevation angle to uncertainties during 0 < t < 20s even without external disturbances and faults. Note that external disturbances and faults appear during time interval 20< t < 60s, where elevation and pitch angle responses are heavily influenced and increase oscillations when system controlled by controller law designed using backstepping method but without any observers" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure5.13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure5.13-1.png", "caption": "Figure 5.13 Angle gauge block combination", "texts": [ " One jaw can be used in conjunction with a length bar combination and the base to form a height gauge. Figure 5.10 Assembled length bars and accessories The two opposite wringing faces are precision D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 5 Standards, measurement and gauging 5 68 3\u00b0 angle gauge is wrung in the opposite direction, with the \u2018minus\u2019 end of the 3\u00b0 over the \u2018plus\u2019 end of the 30\u00b0, this has the effect of subtracting the 3\u00b0 and results in a total angle of 27\u00b0 as shown in Fig. 5.13(b). lapped to a high degree of accuracy of angle and flatness enabling blocks to be wrung together to build up the desired angle. The set shown in Fig. 5.11 comprises: XX five blocks in seconds \u2013 1, 3, 5, 20 and 30. The angle blocks in this set can be wrung together in combination by adding or subtracting to form any angle between 0\u00b0 and 99\u00b0 in 1 second steps. For example, wringing 30\u00b0 and 3\u00b0 angle gauges, as shown in Fig. 5.13(a) with both \u2018plus\u2019 ends together, will result in a total angle of 33\u00b0. If the Setting roller Hinge roller Length Figure 5.14 Sine bar Figure 5.11 Set of angle gauge blocks Figure 5.12 Set-up using angle gauge blocks The main requirements are: (a) the mean diameters of the rollers shall be equal to each other within 0.0025 mm; (b) the upper working surface shall be parallel to the plane tangential to the lower surface of the rollers within 0.002 mm; (c) the distance between the roller axes shall be accurate to within 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001203_s12541-014-0415-9-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001203_s12541-014-0415-9-Figure2-1.png", "caption": "Fig. 2 3D model of test gears", "texts": [ " The heater and cooling coil allow the settling of a constant oil temperature measured by the temperature sensor.7 The test gears which were used for this contact fatigue test were designed and manufactured in accordance with type C test gears of FZG gear test parameters and the relevant provisions of GB/T1422993. In order to fully reflect the carrying capacity of the gear, random sampling was used for the test gears. Test gears are the standard gears and the dimensions are given in Table 1, the gears 3D model which was used for this test is shown in Fig. 2. And they were detected where the tooth profile and tooth root arc line met the test gear requirements. The surface hardness of the gear is 60~62 HRC, the core hardness is 31~40 HRC, the carburizing depth is 0.8~1.0 mm and the residual austenite content is less than 30%. The gear precision grade is 5 (ISO 1328-1: 1995). The Gear material used in this study is alloy steel 20MnCr5. Metallographic samples were prepared according to standard procedure of GB/T4336-2002, the nominal and measured chemical composition of the 20MnCr5 material is given in Table 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002782_0951192x.2020.1815851-Figure27-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002782_0951192x.2020.1815851-Figure27-1.png", "caption": "Figure 27. Feed rate fn and material extrusion rate fm.", "texts": [ " Fanglin, Q. X. Huang, J. Garcia, Y. Chen, C. Tu, B. Benes, H. Zhang, D. Cohen-Or, and B. Chen. 2016. \u201cConnected Fermat Spirals for Layered Fabrication.\u201d In ACM Transactions on Graphics. Vol. 35. Anaheim, CA: Association for Computing Machinery. doi:10.1145/2897824.2925958. This appendix reports how the nozzle feed rate and material extrusion rate are decided (in both hardware and software). Actually, the variation of the extrusion rate fm between any two adjacent NL points (e.g., pi and pi\u00fe1 in Figure 27 below) has a great influence on the amount of extruded material, which means the extrusion system of the printer may not be able to keep the preset extrusion rate at some NL points. For example, the step motor used in the extrusion system may have difficulty in maintaining a precise control of fm. As such, the printing quality of the surfaces would become uncontrollable. Accordingly, a constant extrusion rate fm is preferred in any typical FDM printing process. Conceivably, for any two neighboring NL points, the volume of the extruded material must be synchronized with the re-solidified material (Chen et al. 2019). As illustrated in Figure 27, if the nozzle travels from pi to pi\u00fe1 , the material before melting need to move precisely from ei to ei\u00fe1. Owing to this requirement on volumetric equilibrium and the fact that fm is a constant, the feed rate fn should be regulated by the following equations: 1 4 \u03c0whfn \u00bc \u03c0r2 mfm (22) fn \u00bc 4r2 m wh fm (23) where rm denotes the radius of the filament before melting, and w and h are the major axis length (i.e., the nozzle diameter w \u00bc 2rn) and the minor axis length (i.e., the layer thickness) in the elliptic model, respectively. The nozzle feed rate fn of Equation (23) can be numerically calculated for each sample NL point on the surface. It is worth mentioning that the determination of fn is only related to the layer thickness (h in Equation (23)) at the point (e.g., point pi in Figure 27), indicating that the feed rate fn at every NL point on the surface can be pre-calculated once the constant fm is decided." ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001023_robio.2014.7090430-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001023_robio.2014.7090430-Figure11-1.png", "caption": "Fig. 11. Coordinate system of PHARAD.", "texts": [ " To evaluate the reproduction accuracy, we attached the foot plate to a subject\u2019s foot and compared the reproduced motions with the input motions. In this experiment, is the posture of the foot plate measured by the PHARAD, and is the posture of the foot plate measured by the MCS. We evaluated the measurement performance of the PHARAD by comparing with . Figs.11 and 12 show the coordinate systems of the base frame and foot plate, respectively. The coordinate system of the MCS was matched to the coordinate system of the base frame using the calibration plate, i.e., the black L-shaped plate in Fig. 11. Markers for the MCS were set on the foot plate, as shown in Fig. 12. We used the measured value of marker b as the 3D position of the foot plate. The rotation angle of the foot plate was calculated using a rotation matrix, which rotated the coordinate system of the foot plate. Eleven MCS cameras were used to measuring the foot plate motions. The sampling frequencies of the PHARAD and MCS were set to 100 and 250 Hz, respectively. We conducted two experiments: static and dynamic. Static experiment: The MCS and PHARAD measured the posture of the foot plate, which was fixed for 3 s" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001755_978-3-319-26327-4-Figure4.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001755_978-3-319-26327-4-Figure4.1-1.png", "caption": "Fig. 4.1 Virtual setup and flow diagram for the VSL approach. a Virtual experimental setup for a VSL task consisting of a robot manipulator, a vision sensor, a set of objects, and a workspace. b A high-level flow diagram of VSL illustatrating the demonstration and reproduction phases", "texts": [ " The VSL approach is a goal-based robot learning from demonstration approach. It means that a human tutor should demonstrate a sequence of operations on a set of objects. Each operation consists of a set of actions, for instance a pick action and a place action. In this chapter, we only consider pick-and-place object manipulation tasks in which achieving the goal of the task and retaining the sequence of operations are particularly important. We consider the virtual experimental setup illustrated in Fig. 4.1a which consists of a robot manipulator equipped with a gripper, a tabletop workspace, a set of objects, and a vision sensor. The sensor can bemounted above the workspace to observe the tutor performing the task. Using VSL, the robot learns new skills from the demonstrations by extracting spatial relationships among objects. Afterwards starting from a random initial configuration of the objects, the robot can perform a new sequence of operations which ultimately results in reaching the same goal as the one demonstrated by the tutor. A high-level flow diagram shown in Fig. 4.1b illustrates that VSL consists of two main phases: demonstration and reproduction. In the demonstration phase for each action a set of observations is recorded which is utilized for the match finding process during the reproduction phase. In this section, first, the basic terms for describing VSL are defined. Then the problem statement is described and finally the VSL approach is explained in details. The basic terms that are used to describe VSL consist of: World: the workspace of the robot which is observable by the vision sensor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002976_s10846-020-01288-9-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002976_s10846-020-01288-9-Figure2-1.png", "caption": "Fig. 2 Kinematic bicycle model representing the AMR", "texts": [ " In this section, we present the physics-based model of the AMR to be used for energy-conscientious optimal control and trajectory generation for task completion after scheduling. The overall physical model consists of (a) a kinematic bicycle model of the AMR for trajectory tracking, (b) an actuator model representing the electric drive on the AMR, and (c) a battery model for state-of-charge (SOC) analysis. These models will be explained in detail in this section. For the proposed energy-based optimization, we propose using a kinematic bicycle model, shown in Fig. 2, for modeling the AMR. The set of non-linear equations describing this bicycle model is as follows [42]: X\u0307 = vx cos(\u03c8 + \u03b2) (1) Y\u0307 = vx sin(\u03c8 + \u03b2) (2) \u03c8\u0307 = vx lf sin(\u03b2) (3) v\u0307x = Fx/m (4) \u03b2 = tan\u22121 ( lr lf + lr tan(\u03b4f ) ) (5) where X and Y are the coordinates of the center of mass of the AMR in the inertial (global) frame, vx is the velocity of the AMR, Fx is the force acting on the AMR in the direction of the velocity vector, m is the total mass of the AMR, \u03c8 is the heading angle of the AMR in the global frame, \u03b4f is the steering angle input at the front wheel (assuming a front steered AMR), O is the instantaneous center of rotation, lf and lr are the distances of the front and rear axles respectively from the center of mass, and \u03b2 is the angle between the velocity vector makes and the longitudinal axis of the AMR" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002092_0954406216682768-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002092_0954406216682768-Figure5-1.png", "caption": "Figure 5. Force analysis of contact areas between the ball screw and the nut. 1 master nut; 2 slave nut; 3 preload spacer; 4 ball screw.", "texts": [ " The contact coefficient Kei can be expressed as follows Kei \u00bc Ji mai 3 2 1 u21 E1 \u00fe 1 u22 E2 2( )1=3 1 sin cos \u00f0 \u00de 5=3 X 1=3 i \u00bc 1, 2 \u00f05\u00de The axial contact force of the nut is Fa \u00bc Pz sin cos \u00bc zFbx \u00f06\u00de where z is the number of the working balls in the nut. In the previous study, the contact process for all the balls is the same based on the assumption that the grooves and balls are the same when exerting preloads. Therefore, the axial force is Fp for the master nut and the slave nut when the preload Fp is exerted, which is presented in Figure 5. Under the action of preloads, the elastic deformation occurs in the contact areas between the ball screw and rolling balls, and the areas between the nuts and rolling balls. The axial deformation for nut A and B is the same in value but opposite in direction. The axial deformation p of the nut can be calculated as p \u00bc Ke1 \u00fe Ke2\u00f0 \u00de Fp z 2=3 \u00f07\u00de The axial stiffness of the master nut is equal to the slave nut and can be expressed as kn \u00bc Fp p \u00f08\u00de Despite that the ball screw and the nut are machined together, machining errors of the grooves still cannot be ignored" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002109_iros.2016.7759653-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002109_iros.2016.7759653-Figure4-1.png", "caption": "Fig. 4. Geometric local planning to avoid the object set. (a) Paths are evaluated clockwise (blue) and counter-clockwise (red). (b) Many objects, the push position (blue) is pushed to free space before looking for valid path. Also, a push position (green) for which no path is found.", "texts": [ " 1) Roadmap-based Global Planning Stage: The roadmap, computed as explained in Section III-A, provides global paths through the environment for the agents toward the goal locations as shown in Fig. 3(a). 2) Local Planning Stage: The local planning finds a path that avoids the objects to reach the push position as shown in Fig. 3(b). With potentially many objects in the environment, we have found that geometric local planning performs better than a tree-based search. The object set geometry is explored clockwise and counterclockwise around the object set, keeping the shortest path segment as shown in Fig. 4. This is repeated until the goal location is reached. Additional logic is added to push the start and goal locations to free areas before starting this process if they get in collision. 3) Creating Path from Action: Once all the agents have reached the specified push positions, they are ready to start pushing actions. The agents generate a path from their current location to a position in the direction of the goal and far away from the agent to allow the generation of a force of a maximum magnitude specified by the user" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001601_s00170-016-8487-6-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001601_s00170-016-8487-6-Figure6-1.png", "caption": "Fig. 6 Two distinct illustrations of geometrical parameters of the Exp-ECAE die", "texts": [ " Therefore, the optimization of die geometry was conducted to induce the plastic strain as large and as homogeneous as possible into the sample. To provide an overall view about the optimization process, Fig. 5 depicts the flow chart of the applied approach. Geometrical parameters (as shown in Fig. 1), including sphere diameter (D1), channel diameter (D0), fillet radius at the intersection of the sphere and the channels (r), and eccentricity of the ball cavity center with respect to the centerlines of the channels (e) were nondimensionalized then involved in the Exp-ECAE optimization. These parameters are also presented in Fig. 6a. As this figure shows: 0\u2264R0 \u00fe r \u00fe e\u2264R1 \u00f05\u00de According to Fig. 6a due to unclear position of the fillet radius center, Eq. 5 is a rough description. Dividing this equation by R1 yields: 0\u2264R0 . R1 \u00fe r . R1 \u00fe e . R1\u22641 \u00f06\u00de Dimensionless parameters of \u03b1, \u03b2, and \u03b3 are defined as below: \u03b1 \u00bc R0 . R1 \u03b2 \u00bc r . R1 \u03b3 \u00bc e . R1 8>>< >>: \u00f07\u00de Thus, Eq. 6 could be restated as: 0\u2264\u03b1\u00fe \u03b2 \u00fe \u03b3\u22641 \u00f08\u00de \u03b1, \u03b2, and \u03b3, as dimensionless parameters, are appropriate to be utilized through the optimization process. The die geometry is symmetrical with respect to the bisector of the centerlines of channels. Based on Fig. 6b: 0\u2264R0 \u00fe r\u2212e\u2264R1 ffiffiffi 2 p . 2 \u00f09\u00de Similar to Eq. 5, this equation is also an approximation. Dividing Eq. 9 by R1 results in: 0\u2264R0 . R1 \u00fe r . R1\u2212e . R1\u2264 ffiffiffi 2 p . 2 \u00f010\u00de Substituting Eq. 7 into Eq. 10 yields: 0\u2264\u03b1\u00fe \u03b2\u2212\u03b3\u2264 ffiffiffi 2 p . 2 \u00f011\u00de Equations 8 and 11 introduce the feasible region (FR) for Exp-ECAE examinations. This region is illustrated in Fig. 7a. \u03b1 represents the inversion of expansion ratio (R1/ R0). As this figure shows, \u03b1 can take zero which means applying an infinite expansion ratio" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002775_j.triboint.2020.106669-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002775_j.triboint.2020.106669-Figure8-1.png", "caption": "Fig. 8. The finite element model of twin-discs contact.", "texts": [ " Therefore, according to the vibration characteristics of sliding-rolling contact, the roller speed can be obtained and thus can provide valuable information for skidding detection. The vibration frequency analysis indicates that the dynamic behaviour under sliding-rolling contact is determined by the contact property between the tribo-pair. To explore this contact mechanism, a twodimensional finite element model is constructed and the surface contact stress analysis is performed by using ABAQUS/Explicit. In order to reduce the meshing and calculation time, a simplified finite element model consisting of two \u03c612 mm discs is adopted in this work, as shown in Fig. 8. This simplification is different from that in the experiment. With the purpose of stress analysis, such simplification is reasonable because the contact stress in the current analysis maintains the equivalence with that of the experiment. Besides, the width of the contact area in the simulation model is much smaller than the equivalent radius of curvature [36]. Therefore, the influence of the simplification of C. Xu et al. Tribology International 154 (2021) 106669 the geometry is limited in the current case", " The boundary conditions of the rotation and y-direction displacement are applied to the upper disc, and a normal load is applied to the centre. However, only the rotating boundary condition is applied to the lower disc. Two discs are in friction contact with a friction coefficient of 0.2 and the model is a mesh of CPS4R elements. The material property and contact conditions are the same as those described in Section 2.3. The analysis is focused on the stress distribution of the contact region at the pure-rolling and sliding-rolling instants. Fig. 8 illustrates the contact region of two discs. The lower disc is selected as the research object. The 1\u201313 element nodes, of which the 7-node is positively the center, are marked as tracking targets in the contact region. A pure-rolling motion and a sliding-rolling motion are simulated respectively. The entire analysis process consists of two stages. At the first stage (0\u20130.001 s), a normal load of 50 N is gradually applied and then keeps constant. Referring to the experimental conditions in Table 2, the applied load should be 200 N (the corresponding maximum contact stress is 334 MPa calculated by Hertz contact theory)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002627_0954406220945728-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002627_0954406220945728-Figure7-1.png", "caption": "Figure 7. Analysis of remaining two joint interfaces.", "texts": [ " The analysis of the elastic beam is shown in Figure 6(b), and the deformation at the free end of the elastic beam is zZ \u00bc Gh EZAZ LZ \u00bc g4 \u00f017\u00de xZ \u00bc GhLh LZ \u00fe zZ\u00f0 \u00de 2 cos sin 2EZI Y Z \u00bc g5 cos sin \u00f018\u00de yZ \u00bc GhLh LZ \u00fe zZ\u00f0 \u00de 2 sin 2EZI X Z \u00bc g5 sin \u00f019\u00de XZ \u00bc GhLh LZ \u00fe zZ\u00f0 \u00de sin EZI X Z \u00bc g6 sin \u00f020\u00de YZ \u00bc GhLh LZ \u00fe zZ\u00f0 \u00de cos sin EZI Y Z \u00bc g7 cos sin \u00f021\u00de Similarly, four new kinematic parameters are introduced in the analysis of the joint interface between Band Z-axis: g4, g5, g6, g7\u00bd , and corresponding gravity deformation matrix is Z BTg \u00bc c YZ 0 s YZ xZ 0 1 0 0 s YZ 0 c YZ zZ 0 0 0 1 2 6664 3 7775 1 0 0 0 0 c XZ s XZ yZ 0 s XZ c XZ 0 0 0 0 1 2 6664 3 7775 \u00bc c YZ s XZs Y Z c XZs Y Z xZ 0 c XZ s XZ yZ s YZ s XZc Y Z c XZc Y Z zZ 0 0 0 1 2 6664 3 7775 \u00f022\u00de The analysis of the joint interface between Z- and Y-axis is shown in Figure 7(a), and the analysis of the joint interface between Y- and X-axis is shown in Figure 7(b), where blue lines represent elastic beam deformation models and are fixed at one end and free at the other end. Since the analysis of these two joint interfaces is similar to the process above, it will not be expanded to avoid repetition. Furthermore, it should be noted that the gravity deformation matrix of the joint interface between Z- and Y-axis Y ZTg takes account of the weight of the Z-axis, and introduces six kinematic parameters, while the gravity deformation matrix of the joint interface between Y- and X-axis X YTg takes account of the weight of the Z-axis and the Y-axis, and introduces eight kinematic parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-FigureA.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-FigureA.1-1.png", "caption": "Fig. A.1 Three reflectance maps for different lighting directions", "texts": [ "19 Dependency of the image based translation estimation result of the rotation estimation quality . . . . . . . . . . . . . . . . 93 Figure 5.20 Dependency of the volume based translation estimation result of the focal length . . . . . . . . . . . . . . . . . . . . . . . . . . 93 Figure 5.21 Robot work cell equipped with a photometric stereo based vision sensor . . . . . . . . . . . . . . . . . . . . . . . . . 94 Figure 5.22 Results of the normal map based pose estimation technique on real data . . . . . . . . . . . . . . . . . . . . . . . . . . . 95 Figure A.1 Three reflectance maps for different lighting directions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 105 Figure A.2 Illuminated scene from the camera\u2019s point of view of the single stripe system. . . . . . . . . . . . . . . . . . . 107 Figure A.3 Example of the line voting algorithm . . . . . . . . . . . . . . . . . 108 Figure A.4 Single stripe pattern normal map acquisition principle . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 109 Figure A", "1 can be solved for \u03b1. Thus, using one light source only gives information about one DOF of the surface normal. To completely reject the ambiguities, at least three images have to be taken to solve for the two DOFs (p and q) of the normals. p and q are the partial derivatives of the surface f (x, y): p = \u03b4 f (x, y) \u03b4x (A.3) q = \u03b4 f (x, y) \u03b4y (A.4) The normal n can then be written as n = [p, q,\u22121]T . (A.5) To illustrate this graphically, the reflectance properties can also be represented in reflectance maps (see Fig.A.1). The detected brightness of the surface is dependent of the angle of its normal and the incident light direction. Therefore, in case of identical incident light and viewing direction, circular structures are visible in the Appendix A: Data Acquisition 105 reflectance maps; on each of them the surface has the same brightness. Due to these structures, the described ambiguities arise. Three of these \u201ciso brightness contours\u201d, under different illumination directions, intersect in exactly one point which defines one unique surface normal" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000692_reduas47371.2019.8999700-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000692_reduas47371.2019.8999700-Figure1-1.png", "caption": "Fig. 1. A Tilt-wing system with partially tilting wings in CTOL and VTOL mode", "texts": [ " The tilt-wing flight dynamics, which was investigated for this research project, has been developed by a research group of the Department of Aerospace Engineering of Chungnam National University (CNU), South Korea [11]. The tilt-wing aircraft system is a research prototype which is utilised for novel studies on the system dynamics and the testing of novel control approaches on that aircraft system. The UAV is a small scale tilt-wing aircraft system. Unlike other tilt-wing systems, the wings of that aircraft are just partially tilting, as displayed in figure 1, which means that the inner part of the wing keeps aligned with the fuselage, while the outer part rotates upwards to a tilt angle of 88 degrees. The stall speed in the CTOL mode for the aircraft is 16 m/s, while the cruise speed in that mode is 30 m/s. The aircraft configuration has two lift generating propellers in the front, which both are pivoting together with the tilting part of the wing. For the pitch control of the aircraft during the hover mode, a tail rotor is located on the rear of the aircraft" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001020_s00170-015-7033-2-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001020_s00170-015-7033-2-Figure1-1.png", "caption": "Fig. 1 Comparison of cutting characteristics. a Cutting characteristics of regular torus milling cutter (R=1); b cutting characteristics of regular torus milling cutter (R=2); c cutting characteristics of ball-end milling cutter (R=5); d cutting characteristics of double-circulararc torus milling cutter with double-circular-arc", "texts": [ " The double-circular-arc torus milling cutter designed in this work combines double-circular-arc convex and concave cutting edges to form an equal-helix-angle integral milling cutter. The cutting edge consists of main-circular-arc edge and subcircumferential edge. The milling cutter controls the cutting edge angle through the edge\u2019s radian to reduce the chip\u2019s thickness, which can improve the metal removal rate and reduce the radial force and spindle vibration. The convex cutting edge can eliminate residue stepped surface and ensure cutting uniformity. Figure 1 shows the comparison of cutting characteristics of four different milling cutters under depth of cut ap=0.5 mm. As shown in Fig. 1, the cutting edge angle Kri is: Kri \u00bc 1 2 cos\u22121 R\u2212ap R \u00f01\u00de where R is the radius of the main cutting edge; ap is the depth of cut. The maximum chip thickness Tm is: Tmax \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ap\u22c5 2R\u2212ap2 q R \u22c5Fz \u00f02\u00de The average chip thickness Ti is: T i \u00bc Tmax 2 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ap\u22c5 2R\u2212ap2 q 2R \u22c5Fz \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2R\u2212ap3 4R2 s \u22c5Fz \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi ap 2R \u2212 ap3 4R2 s \u22c5 Fz \u00f03\u00de As ap3 4R2 \u221d0 and sinkri \u00bc ffiffiffiffi ap 2R q , the relationship between the average chip thickness Ti and feed per tooth Fz due to different cutting edge angle can be expressed as: where i is the ith cutting edge. From Fig. 1, the smaller Kri generates smaller average chip thickness Ti. If all of the milling cutters in Fig. 1 have the same feed rate with the new designed double-circular-arc torus milling cutter, the ratio of feed rate is: Fz\u2212i \u00bc Ti T4 \u22c5Fz \u00bc sin Kri sin Kr4\u00f0 \u00de \u22c5Fz \u00f05\u00de Through Eq. (2), it is shown that the double-circular-arc torus milling cutter can increase three times feed rate. The machining time can be reduced even under small depth of cut. As shown in Fig. 1a, regular torus milling cutter suffers high cutting force under high depth of cut, and cutting vibration is caused at high feed rate. The double-circular-arc torus milling cutter has a circular-arc edge involving in cutting and forms small cutting edge angle. Cutting force is decreased under low depth of cut, as the direction of main cutting force is along the spindle and radial cutting force decreases. The requirements to the torque and power of machine tool are reduced. The cutting stability is ensured under high feed rate and cutting vibration will be suppressed", " The forces in the cutting process can be expressed as: F \u00bc Fcsin\u03b30 \u00fe Ftcos\u03b30 \u00f011\u00de N \u00bc Fccos\u03b30\u2212Ftsin\u03b30 \u00f012\u00de Fs \u00bc Fccos\u03c6\u2212Ftsin\u03c6 \u00f013\u00de Fn \u00bc Fcsin\u03c6\u00fe Ftcos\u03c6 \u00f014\u00de where Fc is the cutting force on the direction of cutting speed; Ft is cutting force on the direction of depth of cut; \u03c6 is the shear angle; and \u03b30 is the rake angle. From the above analysis, the shear stress \u03c4s and normal stress \u03c3s remain unchanged in the same cutting condition. The double-circular-arc torus milling cutter can increase the chip width and tool-chip contact length. Therefore the friction stress \u03c4f and normal stress \u03c3f is decreased, which results in a decreasing of the primary stress on the rake face. As shown in Fig. 1d and Eq. (6), the milling cutter has a long effective contact length and the stress in cutting is decreased. The milling cutter\u2019s life can be increased. Compared with regular torus milling cutter, the advantages of double-circular-arc torus milling cutter are summarized as follows: 1. It has small entrance angle and can form thin chip in cutting process. The feed rate can be increased in cutting. 2. Low depth of cut can be used. The direction of main cutting force is along the spindle and cutting vibration can be decreased at high feed rate" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001985_1.4034768-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001985_1.4034768-Figure3-1.png", "caption": "Fig. 3 (a) Stress of rotors, (b) stress of disk b, and (c) stress of axial assembly interfaces at 7500 rpm", "texts": [ " (3) The assembly contact interfaces are also the important components of the static model. The augmented Lagrangian contact algorithm involving friction [9] is used to calculate normal and tangential stress of interfaces. This method is also used in the software ANSYS. (4) The load in this model is pretightening force (188 KN for each rod) and centrifugal force (determined by working speed 7500 rpm). 2.2 Static Results of Defective Rod-Fastened Rotor With Flatness Error. Static analysis is carried out, and Fig. 3 shows that the stress of perfect rod-fastened rotor is distributed uniformly while obvious asymmetrical stress occurs in the defective rotor at the surrounding area of flatness error. As a result, the machining error causes mass eccentricity e, and uneven stress leads to axis-center displacement rb for disk b (Fig. 4). Compared with perfect rotor, the flatness error on defective disk b brings about mass eccentricity (e\u00bc 0.88 lm) and it remains unchanged from beginning to end; rb equals 3.08 lm after pretightening process and it reaches 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001310_1.4029709-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001310_1.4029709-Figure1-1.png", "caption": "Fig. 1 Schematic view of a journal air bearing", "texts": [ " [6] was limited to the case of a point mass rotor. In this paper, the suitability of the proposed approach to deal with more complex geometries is demonstrated through an example of a flexible rotor supported by two air bearings. The dynamic characteristics of air journal bearings are highly affected by the compressibility of the lubricant. For instance, Figs. 2 and 3 show the variation of stiffness and damping coefficients with the excitation frequency of a circular air journal bearing schematically depicted in Fig. 1. The geometric characteristics of the bearing are L \u00bc 38 mm, R \u00bc 19 mm, CR \u00bc 21lm, the rotor is centered and its rotational speed is X \u00bc 25 krpm. The dynamic coefficients were obtained under the small perturbation assumption by solving the compressible fluid Reynolds equation. The flow is supposed isothermal at 293K. Figures 2 and 3 show that the dynamic coefficients are practically constant at low excitation frequencies. As the frequency increases, the bearing becomes stiffer up to an asymptotic value while the damping decreases toward zero" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002587_lra.2020.3007466-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002587_lra.2020.3007466-Figure1-1.png", "caption": "Fig. 1. 3D cad model of the sina slave robot demonstrating (a) coordinate systems and the serial robot, (b) the two DOF agile-eye mechanism, and (c) the tool\u2019s snake-like wrist.", "texts": [ " This robot consists of an RRP serial robot and a parallel agile-eye head mechanism, to manipulate a snake-like surgical instrument. A model-based approach is used to develop a calibrated FK model of the robot with limited experimental data (Section III). Then a huge artificial input-output data is generated in simulation software and used to train an ANN, representing the IK of the robot (section IV). Finally, after implementing the ANN into the robot controller, its performance in improving the tracking accuracy of the robot is assessed (section V). The schematics of the \u201cSina\u201d slave robot is shown in Fig. 1. The robot has seven degrees of freedom (DOFs). The first three DOFs are used to orient (\u03b81 and \u03b82) and insert (d3) the wristed laparoscopic instrument inside the patient\u2019s abdomen, and the other four DOFs are used to manipulate the wrist (\u03b84 and \u03b85), grasp and roll of the instrument. The roll and grasp DOFs can be independently manipulated from the master robot without the need to solve the geometric kinematic transfer between the master and slave robot. Hence. they do not affect the geometric Fig. 1 shows three coordinate systems (CSs). The reference CS, denoted by XBYBZB , is located at the robot\u2019s remote center of motion (RCM), in a way that ZB aligns with the axis of the first DOF (\u03b81) and XB lies in the horizontal plane pointing outward. The second CS, XAEYAEZAE , is located at the end of the serial mechanism and at the center of the 2 DOF parallel agile-eye mechanism. The axis orientations of XAEYAEZAE follows the Denavit-Hartenberg (DH) convention for the three DOF serial mechanism (\u03b81, \u03b82, and d3)", " The first two rows correspond to DH parameters of the serial robot. The active joint offsets are associated with errors in defining the home position of each degree of freedom. Row three contains the link lengths for the non-symmetric agile-eye mechanism, and row four is the active joint offset, i.e., home offset, for agile-eye. Finally, from row five, k is the ratio between tool\u2019s wrist inclination and handle inclination, w is the length of the snake-like wrist, and t is the tooltip length (refer to Fig. 1). The effect of the length of the tool\u2019s rod, i.e. L in Fig. 1(a) is reflected by \u03b4d3 and so it has not been considered separately. Authorized licensed use limited to: Middlesex University. Downloaded on July 18,2020 at 03:54:57 UTC from IEEE Xplore. Restrictions apply. The end-effector pose can be described by the position, PEE , and orientation of the tooltip, ZEE , which can be obtained as PEE = BRAE AERW WPEE and ZEE = BRAE AERW WZEE , (1) where BRAE and AERW are the rotation matrix between the base and the agile-eye CSs, and the agile-eye and the wrist CSs, respectively", " the angle between the normal vector of the tool\u2019s plane and ZAE , and then move it in negative ZAE direction for the amount of L, i.e. the length of the tool\u2019s rod. To obtain \u03b2, we should first find ZH = BRAE \u00d7 AEZH , for which BRAE has been already derived and AEZH can be obtained by solving the FK of the agile-eye mechanism. The FK solution for this mechanism is provided in [27], based on the link parameters of the mechanism (\u03bb, \u03b3, and \u03be) and the active joint parameters (\u03b84 and \u03b85). Now, by knowing theZH ,\u03b2 is calculated as the angle between YAE and (ZAE \u00d7 ZH) (Fig. 1(b)). Finally, by assuming a constant curvature for the snake-like wrist, the expressions for WZEE and WPEE are WZEE = [s\u03c8 0 \u2212 c\u03c8] T and WPEE = [ w \u03c8 (1\u2212 c\u03c8) + ts\u03c8 0 \u2212 l \u2212 w \u03c8 s\u03c8 \u2212 tc\u03c8 ]T , (2) where \u03c8 = \u03c6\u00d7 k. k is the ratio between wrist inclination and handle inclination and \u03c6 is the angle between ZAE and ZH (Fig. 1). Note that, throughout the paper, sx and cx (x being a Greek letter) denote sin(x) and cos(x) and si and ci (i being a number) denote sin(\u03b8i) and cos(\u03b8i), respectively. For FK calibration, we used the system least squares error (LSE) method [12]. A set of markers was attached to the base and another to the end-effector of the robot (Fig. 2). During the calibration tests, the 3D position of the marker sets were captured by an infrared tracker (NDI Hybrid Polaris Spectra, Northern Digital Inc., Canada), with capturing frequency of 60 Hz and overall RMS tracking error of 250\u03bcm", " The results of the extended calibration show that the highest deviation between the DH nominal and calibrated parameters come from the initially zero DH parameters. This was expected, since the spherical links (\u03b11 and\u03b12) were fabricated using highaccuracy 5 axis CNC machining, hence, the errors came mostly from the links\u2019 twist and assembly inaccuracies. IV. INVERSE KINEMATIC MODEL A. Inverse Kinematic Solution for the Calibrated Robot The aim of IK solution is to obtain the input parameters (\u03b81, \u03b82, d3, \u03b84 and \u03b85) for a given tooltip pose (PEE andZEE). From Fig. 1(c), the following geometric equations were derived: w \u03c8 (1\u2212 c\u03c8) + ls\u03c8 = \u2016PEE\u2016s\u03c1 and t+ w \u03c8 s\u03c8 + lc\u03c8 = \u2016PEE\u2016c\u03c1, (3) where \u03c1 is the angle between PEE and ZEE . Solving Eq. (3) results in l and\u03c8. Having the value for l, one can simply calculate d3 asL\u2212 l, whereL is the length of the tool\u2019s rod. AlsoZW was found by rotating ZEE around YW for the amount of \u03c8 + \u03c0, where YW can be described as YW = ZEE \u00d7PEE \u2016ZEE \u00d7PEE\u2016 . (4) Then, ZAE = ZW can be used to solve the IK for the serial spherical mechanism and obtain \u03b81 and \u03b82", " However, for extended calibration model, since all the DH parameters are Authorized licensed use limited to: Middlesex University. Downloaded on July 18,2020 at 03:54:57 UTC from IEEE Xplore. Restrictions apply. non-zero (refer to Table IV), ZAE =\u23a1 \u23a2\u23a3 (. . .)c1c2 + (. . .)c1s2 + (. . .)s1c2 + (. . .)s1s2 + (. . .)s1 (. . .)c1c2 + (. . .)c1s2 + (. . .)s1c2 + (. . .)s1s2 + (. . .)c1 (. . .)c2 + (. . .)s2 + (. . .) \u23a4 \u23a5\u23a6 , (5) where (. . .)s denote non-zero coefficients that depend on\u03b11,\u03b12, \u03b13, and \u03b83. Having \u03b81, \u03b82 and d3, we can obtain YAE and then \u03b2, which is the angle between YAE and YW (Fig. 1(b)). Also knowing that \u03c6 = \u03c8/k, AEZH can be obtained. Having AEZH , one can determine \u03b84 and \u03b85 by solving the IK of the agile-eye mechanism (provided in [27]). Solving the inverse kinematics for the nominal model needs nonlinear numerical method to obtain l and \u03c8 from Eq. (3). The situation is more complicated for the extended calibration model, where the solution for the two DOF serial spherical mechanism (Eq. (5)) also needs resorting to numerical methods. Although a variety of nonlinear numerical methods are available for solving these equations, they involve high computation costs and their convergence is not guaranteed [12], [20], [21]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000052_s00158-019-02201-1-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000052_s00158-019-02201-1-Figure1-1.png", "caption": "Fig. 1 A lubricated periodic surface texture design problem in a rotational tribo-rheometer setting. a Schematic diagram; b simulated periodic sector; c sector design; d full disc design", "texts": [ " Combinations of the fluid models and governing equations are given in Table 1, and are discussed in more detail in Sections 2.1 and 3.1. We compare the results from the viscoelastic models to the Newtonian fluid reference case for the following reasons: first, the simplest models for including viscoelasticity are based on perturbations around the Newtonian fluid model (ordered fluid expansion (Bird et al. 1987)), and second, we are interested in comparing the system performance with viscoelasticity to the conventional Newtonian lubricants. Figure 1 illustrates the design problem presented. We have adapted our previous design optimization strategy (Lee et al. 2017b; Schuh et al. 2017) to design both surface texture topography and non-Newtonian viscometric functions. Figure 1a shows the setup used previously in experiments (Schuh 2015a); the fluid is confined between a flat plate that rotates at a constant angular velocity and a stationary textured surface. A sector shown in Fig. 1b and c is an example design of the surface texture height profiles as a function of r and \u03b8 . Figure 1d shows an example of a fully textured disc using ten periodic sectors. As we extend our study to include nonlinear viscoelastic models, and move from 2-D to 3-D, the computational cost associated with the design problem increases significantly. In our previous study, where we modeled the fluid flow with the (Newtonian) Reynolds equation, the computational cost of the optimization was reduced by using a coarse design mesh that was mapped onto a finer analysis mesh (Lee et al. 2017b). The computational cost of the optimization can also be reduced by linearizing the Reynolds equation (with respect to the design variables) and iteratively solving using a sequential linear programming (SLP) strategy (Lin et al", " We again use this model because of the predicted normal stress generation, which is important in determining the thrust generation with polymer solutions. We limit the parameter ranges to \u03b7pk \u2208 [ 0, 5 2\u03b7s ] , \u03bbk \u2208 [ 10\u22125, 10\u22122 ] (sec), and \u03b1k \u2208 [0.01, 0.5]. The total number of design variables is 3k, where k is the number of relaxation modes in the parameterization. Fluid properties, model parameters, computational mesh resolutions, operating conditions, and design constraint parameters for Cases 0, 1, and 2 in Table 1 are given in this section. The top and bottom discs (gap-controlled rotating disc and fixed textured surface in Fig. 1a\u2013b) have the same outer radius (ro) of 20 mm. The minimum controlled gap height between top and bottom discs (h0) is 269 \u03bcm; this value is used as the lower bound for the texture design gap height variable. The number of periodic sectors needed to construct a full disc (N\u03c6) is 10. The number of mesh nodes for each r -, \u03b8 -, and z-direction is nr = 6, n\u03b8 = 6, and nz = 4, respectively. Note that nz does not apply to Case 1. The angular velocity of the flat plate ( ), as shown in Fig. 1a, is 10 rad/s; solvent viscosity (\u03b7s) and density (\u03c1s) values are 9.624 \u00d7 10\u22123 Pa\u00b7s and 873.4 kg/m3. The number of modes for the Giesekus fluid model is nmode = 2 for Cases 1 and 2. For Case 0, this variable is not defined. The maximum angle for the texture inclination (\u03b8incl = 60\u25e6) is explained in Section 2.2. The design problem considered here is the simultaneous minimization of the input power to the rotating flat plate and maximization of the load-supporting normal force, while constraining the maximum texture inclination angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003356_s11661-021-06148-1-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003356_s11661-021-06148-1-Figure1-1.png", "caption": "Fig. 1\u2014Geometry and dimensions of the compact tensile specimen (in mm).", "texts": [ " To achieve nano-size information (especially carbide characteristics) about the microstructure, a Tecnai G2 F30 S-TWIN field-emission high-resolution scanning transmission electron microscope (HR-STEM) operated at 300 kV and equipped with an EDAX energy-dispersive X-ray spectroscopy (EDS) system was used. Both thin-slice specimens and carbon film replicas with 3 mm diameter were prepared. Moreover, the description of the TEM specimen preparation steps was given in our previous study.[13] Compact tensile (CT) specimens of LAM AerMet100 steel with 37.5 mm 9 36 mm 915 mm size were used for fracture toughness tests. The specific configuration and dimensions of the CT specimens are shown in Figure 1. To facilitate fatigue pre-cracking at a low stress intensity level, the root radius for a straight-through wide slot terminating in a V-notch is 0.1 mm using an electrical discharge machine (EDM). A sharp fatigue pre-crack Table I. Post-LAM Treatment Processes for LAM AerMet100 Steel No. Specimen Designation Preliminary Treatment Subsequent Heat Treatments 1 CG-TB N/A Cryogenic treatment in liquid nitrogen, 2 h, air warming + 482 C, 5 h, air cooling (AC) 2 FEG-TM-CA N/A Special heat treatment 1 + 885 C, 1 h, oil cooling (OC) + cryogenic treatment in liquid nitrogen, 2 h, air warming + 482 C, 5 h, AC 3 CEG-TM 1300 C for 5 hour at 135 MPa, AC Special heat treatment 2 + 885 C, 1 h, OC + cryogenic treatment in liquid nitrogen, 2 h, air warming + 482 C, 5 h, AC 4 FEG-TM Special heat treatment 1 + 885 C, 1 h, OC + cryogenic treatment in liquid nitrogen, 2 h, air warming + 482 C, 5 h, AC 5 FEG-TM-HRA Special heat treatment 1 + 885 C, 1 h, OC + 482 C, 5 h, AC N/A not applicable" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure6.30-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure6.30-1.png", "caption": "Figure 6.30 Micrometer special anvils", "texts": [ " With a digital micrometer, the diameter of the ball or roller can be measured, the micrometer zeroed, and the measurement then made will be the wall thickness shown directly without the need for subtraction. For this type of measurement, an accessory is available called a \u2018ball attachment\u2019, which is a ball (usually 5 mm diameter) held in a rubber boot to hold it safely in place on the micrometer anvil while a measurement is taken. In order to facilitate a wide range of measuring requirements, micrometers are available with a variety of anvils for special purpose applications, a few of which are shown in Fig. 6.30: a) Blade micrometer \u2013 for measuring diameter in a narrow groove (the spindle is non-rotational). b) Tube micrometer \u2013 for measuring the wall thickness of a tube (anvil has a spherical surface). c) Spline micrometer \u2013 for measuring spline-root shaft diameter. d) Disc type micrometer \u2013 for measuring spur and helical gear root tangent. Figure 6.27 Mechanical counter micrometer Figure 6.29 Measuring tube wall thickness (a) Blade micrometer (b) Tube micrometer D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 6 Measuring equipment 6 92 Micrometers with an adjustable measuring force are available for applications requiring a constant low measuring force such as measuring thin wire, paper, plastics or rubber parts which avoids distortion of the workpiece" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002869_access.2020.3036250-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002869_access.2020.3036250-Figure6-1.png", "caption": "FIGURE 6. Fault modes on the gearbox fault dataset by IEEE PHM society.", "texts": [ " As for the signal acquisitions, two accelerometers installed on the input and output shaft retaining plates and a tachometer are used to acquire the vibration data and rotational speed of the input shaft, respectively. The vibration signals of spur gears acquired by the accelerometer on the input side are analyzed in this case. The specific fault modes on this gearbox dataset contain three gear fault modes, three bearing fault modes, the shaft imbalance, the normal gear, and the normal bearing, as illustrated in Fig. 6. In this case study, six different gearbox health states are considered and their fault modes are detailed in Table 1. During the vibration measurement experiments, the rotation speed of the input shaft is 30 Hz and the gearbox operates with a high load. Besides, the sampling rate is 200/3 kHz and the sampling time of each vibration acquisition is 4 seconds. Two different vibration measurements are acquired This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000099_j.matlet.2019.02.095-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000099_j.matlet.2019.02.095-Figure4-1.png", "caption": "Fig. 4. (a) Schematic illustration of the electro-oxidation of glucose by the NPC/Cu(OH)2 hybrid material. (b) The change of current response of the same electrode after repeated experiments.", "texts": [ "The hybrid electrode with high sensitivity was chosen for the next experiment. Fig. 3(e) indicates that a relatively obvious and stable change in current density occurred with the addition of glucose solutions. Fig. 3(f) further shows the change rule between current density and glucose concentration, indicating that the linear response range of hybrid electrode was 10.0 lM\u201310.0 mM or higher with sensitivity of 1.866 mA cm 2 mM 1. Generally, the mechanism of glucose oxidation in alkaline can be summarized as the conversion of Cu (II) and Cu (III). Just as shown in Fig. 4 (a), the initial oxidation of Cu(OH)2 to CuOOH or Cu (OH)4 occurs firstly, then the Cu (III) interacts with the deprotonated glucose in an alkaline solution to form gluconolactone, which hydrolyzes to gluconic acid eventually [20]. The process can be described as: Cu\u00f0OH\u00de2 \u00fe OH ! CuOOH\u00feH2O\u00fe e \u00f01\u00de Cu\u00f0OH\u00de2 \u00fe 2OH ! Cu\u00f0OH\u00de 4 \u00fe e \u00f02\u00de Cu III\u00f0 \u00de \u00fe glucose ! Cu II\u00f0 \u00de \u00fe gluconic acid \u00f03\u00de The electrons generated by the reactions are transported through the nanoporous copper, causing the current signal detected [10]. What\u2019s more, the stability of the electrode was evaluated. Fig. 4(b) shows that no significant change in the current response occurred with the increase of test times, which further proves the potential application of this electrode material in blood glucose detection. The NPC/Cu(OH)2 nanowire array was synthesized on the basis of nanoporous copper through anodic oxidation. Nanoporous copper was prepared via the Zr50Cu50 amorphous precursor. The hybrid electrode showed good electrocatalytic activity and stability in the direction of glucose sensor, whose sensitivity can reach 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001499_978-3-319-06331-7_19-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001499_978-3-319-06331-7_19-Figure1-1.png", "caption": "Fig. 1 Wind turbine drivetrain components (Oyague 2009)", "texts": [ " One method is based on the classical Weibull approach and the other on application of a LogNormal distribution as done e.g. for the fatigue life of welded steel details. Currently, most operating wind turbines use a modular configuration (Hau 2006). Typically, all individual components of the drivetrain are mounted onto a bedplate. The basic drivetrain components are the main bearing, shaft, gearbox, brakes, high-speed shaft and the generator, see Hau (2006) and Lindley (1976). A typical configuration of these components in the nacelle of a wind turbine is shown in Fig. 1. Reliability of wind turbine gearboxes is studied in a number of research projects, e.g. the GRC project at National Renewable Energy Laboratory (NREL), (Oyague 2009). This include as important areas research on fault diagnosis and condition monitoring. Several methods have been considered, such as vibration and acoustic emission (Soua et al. 2013) and Local mean decomposition (Liu et al. 2012). Some studies on probabilistic modeling of failures in wind turbine drivetrain components have been carried out (Dong et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003599_j.ast.2021.106823-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003599_j.ast.2021.106823-Figure1-1.png", "caption": "Fig. 1. A schematic of F-18 model (in flat-spin) integrated with deployable fin.", "texts": [ " Aerodynamic forces and moments are generated by relative airspeed, and it can be altered by disturbing the airflow for controlling the motion of the aircraft. In this section, the dynamics of a deployable vertical fin is incorporated in standard aircraft equations of motion to study its effect in aircraft spin motion and recovery. Also, later in this section, fin sizing and placement have been designed along with the formulation of optimal fin deflection. All the dynamic equations used in this work are written in a body-fixed reference frame [41,42]. A rough schematic of deployable fin integrated with standard F-18 HARV model is shown in Fig. 1. The geometric characteristics and aerodynamic derivatives of this model are shown in Table 1 and Table 2, respectively. The following assumptions are made for problem formulation which are: 1) Aircraft is rigid with vertical plane of symmetry, 2) Moment of inertia is not varying due to fuel consumption, 3) Unsteady aerodynamics, as well as solid angle effect of fin, are not considered and 4) Wind, gust effect also has not been taken into account. Let, = Ixx Izz \u2212 I2 xz , 1 = \u22121 ( Ixz ( Ixx \u2212 I yy + Izz )) , 2 = \u22121 ( Izz ( Izz \u2212 I yy ) + I2 xz ) , 3 = \u22121 Izz , 4 = \u22121 Ixz , 5 = \u22121 yy (Izz \u2212 Ixx), 6 = \u22121 yy Ixz , 7 = \u22121 ( Ixx ( Ixx \u2212 I yy ) + I2 xz ) and 8 = \u22121 Ixx ", " (1) and the angular acceleration in roll axis ( p\u0307 f ) and yaw axis (r\u0307 f ) in Eq. (3), induced by fin can be modeled as p\u0307 f = \u03c1V 2 S f (CL\u03b1) f ( i f + \u03b2 ) z f 2Ixx r\u0307 f = \u2212\u03c1V 2 S f (CL\u03b1) f ( i f + \u03b2 ) x f 2Izz D f = 1 2 \u03c1V 2 S f C D f \u23ab\u23aa\u23aa\u23aa\u23aa\u23aa\u23ac \u23aa\u23aa\u23aa\u23aa\u23aa\u23ad (5) where C D f = C D0 f + k f ( (CL\u03b1) f (\u2223\u2223i f + \u03b2 \u2223\u2223))\ufe38 \ufe37\ufe37 \ufe38 Induced drag coef f icient 2 + C Di f \u2223\u2223i f \u2223\u2223 Here, i f > 0 and i f < 0 has been considered for left and right spin cases, respectively where i f is being measured with respect to plane of symmetry, as shown in Fig. 1. It is important to note here that the lift produced by fin is not producing any drag as there are insignificant vortices being generated, making induced drag negligible. So, the total effective drag produced by fin will be a combination of parasite drag and drag due to fin deflection. In view of the fact that the sideslip angle is small during aircraft flat-spin, hence, drag along body y-axis is considered negligible. The symbols used to denote stability and control derivatives in Table 2 have their usual meaning and elevator actuation is consid- ered as a deflection of left and right elevators in the same sense (\u03b4e = \u03b4el = \u03b4er ). The fin made of a flat plate is attached with the fuselage as shown in Fig. 1 with a setting angle i f in such a way that the airflow during spin first passes over a leading edge and then leaves from trailing edge of the fin. This setup will be favourable on both left and right spin cases, which is an advantage when compared with fin made of a cambered airfoil section. Low thickness to chord ratio of fin (9%) has been considered, which helps in inserting the fin inside fuselage when not in use. It can be observed that lateral and directional motion will be influenced by adding the fin dynamics and it is quite evident that exposure of fin created additional drag" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001868_amm.842.251-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001868_amm.842.251-Figure1-1.png", "caption": "Fig. 1 Tilt-rotor Reference Frames", "texts": [ " In flight dynamics analysis, the longitudinal plane of motion is described by a plane spanned by x- and z-axes with respect to aircraft body. All of the reference frames are originated at the aircraft\u2019s centre of gravity. There are three reference frames which are commonly used to describe the attitude and motion of aircrafts, namely the Body-fixed reference frame, the Wind reference frame, and the Local horizon refence frame [6]. In longitudinal plane, the relative orientation between the three reference axes are represented by angle of attack (\u03b1), pitch (\u03b8), and trajectory angle (\u03b3). Fig.1 illustrates the reference frames. Equation of Motion. The formulation of the tiltrotor is derived from the equations of motion of a rigid body involving some forces and moments working on the center of mass of the vehicle. With some simplifications taken into account, the resulting equation of motion along the body axes can be expressed as follows [3]: = + + sin . (1a) = \u2212 \u2212 cos . (1b) = . (1c) = . (1d) The term X and Z represents the external forces along the x and z axes, while M is the pitch moment, and q is the pitch rate of the aircraft" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002374_j.ymssp.2020.106844-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002374_j.ymssp.2020.106844-Figure3-1.png", "caption": "Fig. 3. The layout of the accelerometers on the seat (a) on a high-speed train (b) schematic diagram.", "texts": [ " The measurement was carried out in a carriage in which the rear two-thirds is a first-class chamber while the front one-third is a VIP chamber. Vibration measurements were made with two double-unit seats at different positions of the first-class chamber, one near the center of the carriage (the front seat) and the other close to the end (the rear seat) (Fig. 2). The seats consisted of a backrest and a seat pan (almost horizontal). The angle of the backrest could be adjusted, however, was kept vertical during the measurement (Fig. 3). During the measurement, two seating conditions were considered, one with the seat seated with one subject and the other occupied by two subjects. The measurement for the two seating conditions was made with the same male subject (test subject for whom the vibration was measured) sitting on the left. The information of this test subject and the subjects sitting beside were listed in Table 1. During the measurement, all the subjects sat in a comfortable upright posture with the back in contact with the backrest and hands resting on the thighs. Note that in this paper, \u2018left\u2019 and \u2018right\u2019 are defined from the perspective of the subject sitting on the seat. A total of 11 channels of accelerations were recorded, as shown in Fig. 3(b). Two SIT-pads conforming to ISO 10326-1 with a sensitivity of 10 mV/ms 2 in every axis were positioned at the seat pan and backrest to measure the vertical and lateral accelerations at these positions. The SIT-pad on the seat pan was positioned beneath the ischial tuberosity of the test subject. The SIT-pad on the backrest was centrally located zd \u00bc450 mm above the seat pan surface. For calculating the roll motion of the seat pan, two single-axis accelerometers from B&W Tech with sensitivities of 10 mV/ms 2 were mounted on the left and right armrests to measure their vertical accelerations", " There were two kinds of excitations, and the first one was one of the ten measurements of the vertical, lateral and roll accelerations measured at the seat base of the front and rear seats on the train, respectively, conforming to the spectrum in Fig. 4. The second one was a tri-axial random acceleration signal defined in the range from 0.5 to 50 Hz. The root-mean-square (r.m.s.) values of the random signal were 0.5 m/ s2 in the lateral direction, 0.5 m/s2 in the vertical direction and 0.75 rad/s2 in the roll direction, and signals in different directions were almost incoherent with each other. The measurement method was the same as that on the train (Fig. 3), and all the data were recorded by HVlab data acquisition system at 512 samples per second via anti-aliasing filter set at 100 Hz. The experiment was approved by Human Experimentation Safety and Ethics Committee of the ISVR at the University of Southampton. For the on-site measurement, the low multiple coherence at some frequencies can be caused by the low energy of the inputs that is not enough to excite the vibration, or the dominating noise in the response. In the laboratory measurement, the latter can be reduced to some degree" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002637_edm49804.2020.9153527-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002637_edm49804.2020.9153527-Figure1-1.png", "caption": "Fig 1. The position PMSM drive structure in the d q\u2212 rotor reference frame.", "texts": [ " Hence, the cogT is independent of the stator currents and can reach up to 4 % of the nominal operating condition [2]. III. VECTOR CONTROL OF POSITION PMSM DRIVE The vector control strategy is realized on the basis of a PMSM dynamic model introduced in the previous section with assuming impressed stator currents, i.e. the PMSM stator windings are supplied from a current-controlled IGBT-inverter with a high switching frequency. The general configuration of PMSM drive with a rotor position vector control for a precision tracking application is shown in Fig.1. This schematic diagram consists of the following main components: \u2013 a vector control module with r\u03c8 orientation and the instantaneous stator currents decomposition into two components: direct (flux-producing) and quadrature (torque-producing) currents; \u2013 a frequency converter for adjusting the electric power flow to the motor; \u2013 the PMSM connected to a mechanical load. As it can be seen from the (1), the high torque sensitivity and the efficient operation of the PMSM drive is achieved, where the direct stator current sdi is set to zero", " 476 Authorized licensed use limited to: Cornell University Library. Downloaded on September 08,2020 at 04:41:27 UTC from IEEE Xplore. Restrictions apply. According to the basic principles of a cascaded control theory, the three-level vector control system of position the PMSM drive contains the inner closed loops of direct and quadrature stator currents and the two outer feedback loops of the rotor speed and position. The rotor angle r\u03b8 is sensed by absolute angle encoder and the rotor speed r\u03c9 is obtained by differentiation. As it can be clear from Fig. 1, the difference between the rotor angle reference value * r\u03b8 and the measured rotor angle r\u03b8 serves as an input signal to the PI-controller. Such structure of a position controller suppresses the error caused by a load disturbance. The speed loop feeds the inner stator current one and produces the reference * sqi for generation of motoring or regenerating torque as a function of the speed error. In such operating condition the PI-structure for the speed controller is most appropriate. As it is mentioned before, the current source in the stator circuits is a basic approach of the PMSM vector control implementing for a high-performance torque production" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002526_tmag.2020.3003268-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002526_tmag.2020.3003268-Figure1-1.png", "caption": "Fig. 1. (a) Magnetic flux lines of the STPMM and (b) prototype of the rotor.", "texts": [ " Therefore, an STPMM can produce a larger amount of magnetic flux than other PM machines for the same PM volume and can achieve a highly competitive cost [1]-[4]. For example, an STPMM that utilizes low-cost ferrite PMs can replace surface-mounted machines with expensive rare-earth PMs [4]. However, the STPMM suffers from increased axial flux leakage around its PMs to the same degree that it concentrates flux [5]-[8]. A part of the flux leaks inside the rotor core because the PMs are arranged symmetrically as shown in Fig. 1. Fig. 2(a) shows the axial magnetic flux density distributions through a three-dimensional (3D) analysis. The flux leaks around the entire rotor core where the PM is inserted as well as in the air gap. The back electromotive forces (EMFs) from the two-dimensional (2D) and 3D analyses of the STPMM indicate a difference of approximately 10% as shown in Fig. 2(b). This EMF decrease is directly related to the increase of input current and copper loss due to the constant torque reduction. Therefore, in the design of an STPMM, it is necessary to consider the axial leakage flux using a 3D analysis model" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000514_j.mechmachtheory.2019.103667-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000514_j.mechmachtheory.2019.103667-Figure4-1.png", "caption": "Fig. 4. Generation of face gears by a shaper considering: (a) the standard configuration without errors of alignment and (b) the compensation of errors of alignment due to shaft deflections.", "texts": [ " The axial displacement of the pinion is given by A1 = OkP1 \u2212 O f P\u2032 1 (15) where P\u2032 1 corresponds to the unloaded position of point P1 in coordinate system Sf. (4) Axial displacement of the face-gear, A2. The axial displacement of the face gear is given by A2 = OlW1 \u2212 O f W \u2032 1 (16) where W \u2032 1 corresponds to the unloaded position of point W1 in coordinate system Sf. The relative errors of alignment due to shaft deflections will be considered during the process of generation of the face gear. Fig. 4(a) shows the generation of a face gear by a shaper in standard configuration without the consideration of errors of alignment. Fig. 4(b) shows the nonstandard configuration in which errors of alignment due to shaft deflections are compensated. Here, only the shaft angle error, \u03b3 , the minimum distance between the pinion and face-gear axes of rotation, E, and the axial displacement of the face gear, A , will be considered for compensation. The pinion axial displacement 2 A1 will have no influence on the generated geometry of the face gear because the pinion and the shaper are cylindrical spur gears. In the standard configuration, the axis of the shaper cutter zs intersects the axis of the face gear z2 at point Of. The shaft angle is \u03b3 (see Fig. 4(a)). In the nonstandard configuration for compensation of errors of alignment due to shaft deflections, the shaft angle becomes \u03b3 + \u03b3 , and there will be a shortest distance between the axes of the gears denoted as E. The axial displacement of the face gear is denoted as A2 and it will be positive if it moves towards the positive direction of axis z2 (see Fig. 4(b)). We suppose that the generating surfaces of the shaper are known and given as rs = rs(us, \u03b8s) (17) where us and \u03b8 s are the surface parameters along the profile and longitudinal directions, respectively. Details on the determination of the generating surfaces of a shaper can be found in [22]. The face-gear tooth surfaces are determined as the envelope to the family of shaper tooth surfaces. The following coordinate systems are applied for generation: (i) movable coordinate systems Ss and S2 rigidly connected to the shaper and the face-gear, respectively, (ii) fixed coordinate systems Sj and Sm, whose origins coincide with Os and O2, respectively, and (iii) auxiliary coordinate systems Sk and Sl, as shown in Fig. 4. During the generation, the shaper and the face-gear perform rotations about axes zj and zm given by angles \u03c8 s and \u03c82, respectively. These rotations are related by \u03c8s \u03c82 = N2 Ns (18) Here, N2 represents the number of teeth of the face gear and Ns the number of teeth of the shaper. The family of shaper surfaces s is represented in coordinate system S2 by r2(us, \u03b8s,\u03c8s) = M2mMmlMl jM js(\u03c8s)rs(us, \u03b8s) (19) where M js = \u23a1 \u23a2\u23a3 cos \u03c8s sin \u03c8s 0 0 \u2212 sin \u03c8s cos \u03c8s 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a6 (20) Ml j = \u23a1 \u23a2\u23a3 1 0 0 E 0 1 0 0 0 0 1 OjOk 0 0 0 1 \u23a4 \u23a5\u23a6 (21) Mml = \u23a1 \u23a2\u23a3 1 0 0 0 0 cos(\u03b3 + \u03b3 ) \u2212 sin(\u03b3 + \u03b3 ) 0 0 sin(\u03b3 + \u03b3 ) cos(\u03b3 + \u03b3 ) \u2212OlOm 0 0 0 1 \u23a4 \u23a5\u23a6 (22) M2m = \u23a1 \u23a2\u23a3 cos \u03c82 sin \u03c82 0 0 \u2212 sin \u03c82 cos \u03c82 0 0 0 0 1 0 0 0 0 1 \u23a4 \u23a5\u23a6 (23) Angle \u03b3 + \u03b3 denotes the shaft angle and it is the angle formed by the axes of the shaper and the face gear due to shaft deflections. When \u03b3 = 90\u25e6 and \u03b3 = 0\u25e6, distance O jOk between the origins of coordinate systems Sj and Sk is equal to the pitch radius of the face gear, rp2 . However, when the shaft angle is not equal to 90\u25e6, as for example when \u03b3 = 90\u25e6 and \u03b3 = 0 in the configuration for compensation of errors of alignment (Fig. 4(b)), distance O jOk is determined as OjOk = rp2 sin(\u03b3 + \u03b3 ) + rps tan(\u03b3 + \u03b3 ) + A1 (24) Here, rps is the pitch radius of the shaper. Similarly, distance OlOm between the origins of coordinate systems Sl and Sm is given by OlOm = rps sin(\u03b3 + \u03b3 ) + A2 (25) All errors of alignment to be compensated are considered in Eqs. (20)\u2013(23). The equation of meshing is determined as fs2(us, \u03b8s,\u03c8s) = ( \u2202r2 \u2202us \u00d7 \u2202r2 \u2202\u03b8s ) \u00b7 \u2202r2 \u2202\u03c8s = 0 (26) The modified surface of the face-gear with errors of alignment compensated during generation is determined in three- parameter form by vector equation r (us, \u03b8 s, \u03c8 s) and equation of meshing f = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003398_s00466-020-01960-9-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003398_s00466-020-01960-9-Figure8-1.png", "caption": "Fig. 8 Residual stress fields for the same additive buildup as Fig. 7 using the part-scale model [32]. Results shown correspond to a timestep 1, b timestep 50, c timestep 250, and d timestep 1000", "texts": [ " In the flat plate example, the high-fidelity multiphysics workflow shows a divot forming and moving as the laser scans; a distinct feature not captured by the part-scale process model. Through the inclusion of additional physics associated with laser interactions, and gas/fluid effects, the resulting cross sectional profile of the high-fidelity coupled workflow is able to much more closely match the experimental observation compared with the part-scale workflow. We performed similar comparisons for the 4-pass LENS additive build simulation. Figure 7 demonstrates the high fidelity thermal/fluid results along with the associated residual stress predictions. Figure 8 demonstrates the residual stresses predicted by the part-scale model for a similar build, again demonstrating that the multiphysics workflow yields a significantly higher residual stress prediction when compared to the part-scale model. Line plots comparing von Mises stress results for the part-scale model and the highfidelity multiphysics workflow at the build/plate interface along the center of the build are shown in Fig. 10. Corresponding temperature line plots are presented in Fig. 9 .The fluid mechanics portion of the workflow was completed in 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002076_j.acme.2016.10.004-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002076_j.acme.2016.10.004-Figure3-1.png", "caption": "Fig. 3 \u2013 Structural diagram of the friction damper elements with spherical friction surfaces: (1) outer spherical rings, (2) inner spherical ring [5].", "texts": [ " For this reason all of the remaining structure of the robot was omitted, because the possibility of its use to install the frictional damper were negligible. A cast-iron block allowed rigid fixing of the lance, and its compact structure provided no additional feedback, which could affect the dynamic behaviour of the lance with attached damper. One of the main tasks of this paper was to design and construct a friction damper, which uses the difference of rotation angles of cross-sections of the lance subjected to bending [5]. The design of the friction damper based on spherical friction surfaces was proposed and executed. Fig. 3 shows a structural diagram of the elements of the damper developed. The friction damper consists of three spherical rings: two outer and one inner ring. It was made of 40HM steel. This steel is intended for machine parts of very high strength and ductility, and parts exposed to varying load, for example axles, crankshafts, discs, rotors, levers, push rods. The final properties of the damper contact layers were obtained by lapping of friction surfaces. The first was abrasive compound with a grain size 6 mm and next one with grain 3 mm" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000897_iccep.2015.7177539-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000897_iccep.2015.7177539-Figure3-1.png", "caption": "Fig. 3. Cross section of the IPMSM structure.", "texts": [], "surrounding_texts": [ "In order to achieve the goals described in the previous sections, the motor has been driven with the following working conditions: \u2022 \u03c9ref range:500\u00f74000 rpm, with steps of 500 rpm (eight different speed conditions); \u2022 load range: 0%\u00f7 100% of rated load, with steps of 25% (5 different load conditions). Therefore, 40 overall working conditions with defined speed and load have been tested. In addition, for each condition, the idref value has been varied from -2.4 A to +2.4 A with steps of 0.2 A, in order to set a specified magnetization level of the motor. Thus, for each test, 25 different magnetization conditions have been realized, obtaining 1000 different overall working conditions and over 5000 sets of measurements. The \u0394P characteristic as a function of idref , parametrized for the proposed load conditions and at 2000 rpm, is plotted in fig.7. As it is possible to notice, the minimum value of power losses is obtained for a negative value of the id current. By changing the applied load, the minimum value of \u0394P slides through higher negative values of the direct-axis current. Figures 8 and 9 show the \u0394P vs idref characteristics parametrized as function of the reference speed, maintaining the applied load fixed at 25% and 75% of its rated value, respectively. As expected, it can be noticed that the minimum values of power losses are obtained for negative values of idref and, by increasing the applied load, the related peaks are detected for higher negative values of idref . These results can be discussed also from another point of view, as shown in fig. 10, which reports the IPMSM efficiency parametrized for different percentages of the rated speed, as function of idref with an applied load fixed at 50%. As expected, it can be noticed that the maximum efficiency is obtained for negative values of idref . In addition, by increasing the speed, the related peak is detected for higher negative values of idref . The minimum efficiency peak is equal to 83% and it is detected at 1000 rpm, while the remaining four efficiency peaks are relatively close to each other, ranging from 88% to 92%. In terms of maximum efficiency, the IPMSM should work between a quarter to half of its rated load. As a matter of fact, for almost every reference speed, in the load interval between 25% and 50% the maximum efficiency value is close to 90%, while outside of this interval the maximum efficiency decreases significantly with increased values of the applied load. This fact can be observed in the surface graph of fig.11, which shows the maximum efficiency values as function of both rated speed and applied load. Therefore, it can be stated that, for each operating condition of the motor, it is always possible to determine a specified idref value that maximizes the IPMSM efficiency, without decreasing the dynamic performances of the drive. Furthermore, this system is suitable for several typologies of brushless motors and their related applications because of its high flexibility." ] }, { "image_filename": "designv11_22_0000698_icpedc47771.2019.9036558-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000698_icpedc47771.2019.9036558-Figure5-1.png", "caption": "Fig. 5. Altitude Controller Block", "texts": [ " The estimated data will be fed to flight control system which in turn produces the desired motor commands. A. Flight Control System Flight control block (FCB) is the heart of the Quadcopter system. The data (waypoints) given from the spreadsheet will be the input to the FCB. 2nd International Conference on Power and Embedded Drive Control (ICPEDC) 117 Authorized licensed use limited to: Fondren Library Rice University. Downloaded on May 18,2020 at 11:21:07 UTC from IEEE Xplore. Restrictions apply. Fig. 5. is the representative block diagram of the fight control block of the Simulink example model of the quadcopter [9]. The reference values are the desired input values. This block controls the yaw, roll, pitch and X - Y - Z positions. Estimated values are sent from the state estimator which determines the previous position and its respective velocity and acceleration values. X-Y block sets the reference values for roll and pitch to attitude block which further gives Tau-Roll, Tau-Pitch commands. In this manner yaw block gives Tau-Yaw and altitude block gives thrust to the control mixer" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002953_0954406220974052-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002953_0954406220974052-Figure2-1.png", "caption": "Figure 2. The simplified scheme of the parallel mechanism of HFFD-6.", "texts": [ " As shown in Figure 1, the parallel mechanism is a modified Delta parallel mechanism with 3- DOF in translational directions. While the serial mechanism has 3 rotational DOF. The parallel and serial structures can be decoupled; thus, the kinematic analysis can be obtained respectively. From Figures 1 and, 2 the parallel mechanism is a horizontally-mounted modified Delta mechanism. It composes a fixed base, a moving platform, three active linkages, and three passive parallelogram linkages. Three identical motors are mounted at the back of the fixed base. As shown in Figure 2, the world coordinate frame Of g of the parallel mechanism is attached to the center of the fixed base. Ai, Bi and Ci represent the connecting points of the fixed base and active linkage, active linkage and passive linkage, passive linkage and moving platform, respectively. The X-axis is parallel to A2A3, Y-axis through the point A1 and Z-axis is perpendicular to the fixed base. Similarly, the coordinate frame fO0g is established. The geometric analysis method is adopted to establish the forward kinematics formula of the parallel mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001221_esda2014-20001-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001221_esda2014-20001-Figure5-1.png", "caption": "FIGURE 5. IN-SITU LASER CLADDING DEVICE\u2019S GEARING TRANSMISSION", "texts": [ " The device includes two servo motors, the first of which is installed in the first carriage and operatively connected to the laser nozzle to control its pivoting angle (see Fig. 4). A second servo motor is installed on the second carriage and is operatively connected to one of the two lower rods by means of a gearing transmission to control the laser nozzle\u2019s longitudinal position. The device\u2019s gearing transmission comprises a spur gear connected to the second servo motor and a toothed rack, which is formed directly by the lower rod. The first servo motor is connected to the laser nozzle through a bush (see Fig. 5). To ensure the device\u2019s positioning and controlled rotation around the crankshaft journal, it comprises the two aforementioned guide-ways and two opposite-guide-ways. Two supporting plates are permanently fixed on the opposite guideways. Furthermore, these supporting plates are connected to each other by two opposite-rods by means of which both perpendicular guide-ways are in fixed connection to each other. When installed on the crankshaft journal, the guide-ways and opposite guide-ways are connected and fixed to each other by means of four adjustable arms" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002730_cnm.3400-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002730_cnm.3400-Figure4-1.png", "caption": "Fig. 4.(a) Deviation diagram between the actual and equivalent length of the strut. (b) The closed vector chain", "texts": [ " The length L1(L3) is the distance from the center of revolute joint R1(R2) to the point B1(B3). The length L2(L4) is the distance from the pointA1(A2) to B1(B4). The length L5 is the distance from the center of revolute joint R3 to the the center of revolute joint R4. The length li(i=1,2,3,4) is the distance from the point Ai (i=1,2) to the point Bi (i=1,2,...,4). The length l5 is the distance from the center of revolute joint R3 to the the center of revolute joint R4. There is a deviation between the length of the actual strut and the equivalent strut, as shown in Fig. 4 (a). The conversion relationship can be expressed as Eq. (1). -=L l m (1) This article is protected by copyright. All rights reserved. where L denotes the actual strut length of R1B1, l is the equivalent strut length of A1B1, and m indicates the length of R1A1. Assume that \u03b1, \u03b2 and \u03b3 are the angles of rotation of the moving coordinate system O'-x'y'z' relative to the fixed reference frame O-xyz around the x-, y-, and z-axes, respectively. Using the z-y-x Euler angle representation method, if the above rotation angles are known, the homogeneous transformation matrix of the moving coordinate system relative to the fixed coordinate system can be expressed as Eq. (2). ( ) ( ) ( ) 1 4 7 2 5 8 3 6 9 = , , , = O O z y x c c c s s s c c s c s s s c s s s c c s s c c s s c s c c r r r r r r r r r \u03b3 \u03b2 \u03b1 \u03b3 \u03b2 \u03b3 \u03b2 \u03b1 \u03b3 \u03b1 \u03b3 \u03b2 \u03b1 \u03b3 \u03b1 \u03b3 \u03b2 \u03b3 \u03b2 \u03b1 \u03b3 \u03b1 \u03b3 \u03b2 \u03b1 \u03b3 \u03b1 \u03b2 \u03b2 \u03b1 \u03b2 \u03b1 \u2032 \u2212 + = + \u2212 \u2212 R R R R (2) where c\u03b1 and s\u03b1 represent cos(\u03b1) and sin(\u03b1), respectively. The closed-loop method is used to solve the inverse kinematics solution of the mechanism. The vector loop is shown in Fig. 4 (b). Bi (i=1,2,3,4) is the position vector of point Bi relative to point O in the fixed reference frame O-xyz, which can be expressed as Eq. (3). +O i O i\u2032 \u2032 \u2032=B R B O (3) Where [ ] ' ' ' = 0 T ix iy iz T b b b y z = \u2032 ' iB O O' is the position vector of point O' relative to point O in the fixed reference frame O-xyz, and B'i is the position vector of B'i relative to the O' in the moving coordinate system O'-x'y'z'. The strut vector li(i=1,2...,5) can be expressed as Eq. (4) in the fixed reference frame O-xyz" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001601_s00170-016-8487-6-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001601_s00170-016-8487-6-Figure14-1.png", "caption": "Fig. 14 Effective strain distributions in the Exp-ECAE product a optimized based on Eq. 12, b optimized based on Eq. 15, and c obtained experimentally", "texts": [ " 12, the GA finally proposed values of 0.450, 0.000, and 0.100 for \u03b1, \u03b2, and \u03b3, respectively. This solution corresponds to the fifth examination of Table 2. Considering Fig. 7b, one can find out that this point of design space lays on the FR border. Based on the FE results, this optimum design induces plastic strain with the average of 3.886 and with the variance of 0.066, which were respectively estimated with 6.83 and 8.82 % errors via ANN. The contour plot of effective strain caused by this optimum design is plotted in Fig. 14a which can be compared with that obtained by means of the manufactured die (Fig. 14c). By reducing \u03b1, studying the material flow reveals that in addition to increase in the expansion ratio, another mechanism intensifies imposed plastic strain as well. For small expansion ratios, the material continuously flows through the whole cavity then moves to the exit channel. By increasing this ratio more than a particular limit (reduction in \u03b1), the flow pattern alters within the ball-shaped hollow. Assuming a constant diameter for the channels, rising the expansion ratio increases the interaction areas between the material and the walls of cavity and, consequently, the corresponding interfacial friction force amplifies", " Reduction in wasted material was another consideration taken into account during the optimization process. In this regard, Eq. 15 was utilized as the objective function, the GA converged to the values 0.751, 0.000, and 0.100 for \u03b1, \u03b2, and \u03b3, respectively. By conducted FE simulation based on these geometrical optimum parameters, \u03b5ave and \u03c32 were measured to be 1.774 and 0.086 which were estimated by GAwith 0.95 and 5.90 % errors, respectively. For this optimum design, the contour plot of effective strain distribution is exhibited in Fig. 14b. As this figure indicates, compared with the first design (the geometry optimized based on Eq. 12), this optimum geometry leads to respectively 46% reduction and 30% increase in \u03b5ave and \u03c32, although assuming a constant radius (R1) for the sphere, the lost material was decreased by a factor of 2.8. Therefore, during the Exp-ECAE the mass loss could be controlled by adjusting \u03b7 in Eq. 15. As the evidence show, various product qualities are achievable by means of the Exp-ECAE operation. This capacity is specifically comparable with the ECAE process, in which the imposed plastic strain is constrained within a narrow range [3] though according to Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003847_j.mset.2021.08.005-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003847_j.mset.2021.08.005-Figure5-1.png", "caption": "Fig. 5. (a) Equivalent Shear Elastic Strain distribution in workpiece (b) Equivalent Shear Elastic Strain VS Time Curve.", "texts": [ " 4 (a) demonstrates the Effective Equivalent Total deformation distribution in the workpiece during the extrusion process and 4 (b) demonstrates the Equivalent Total deformation vs Time curve during simulation in the workpiece during extrusion. The zone in which the total deformation occurs inside the corner region of the channel is known as the main deformation zone. It can be detected that the maximum value achieved for the effective equivalent Total deformation alters from roughly zero to 0.0219 m. Fig. 5 (a) demonstrates the Equivalent Shear Elastic Strain in the workpiece during extrusion and 5 (b) demonstrates the Equivalent Shear Elastic Strain vs Time Curve for the workpiece during simulation. It can be detected from Fig. 5 (a) that the maximum value for the shear elastic strain achieved during extrusion ranges from 0.475 to 0.538 m. xtrusion (b) Equivalent Total Distribution VS Time Curve. In this paper, the Ti-6Al-4V alloy and its associated properties employing simulation analysis with the help of available ANSYS software were observed. The simulation was done by forming a billet of Ti-6Al-4V alloy that was passed through the channel inside the die under the ECAP process. This paper also analyzed several factors and their influence on the workpiece while undergoing ECAP" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003796_s10846-021-01450-x-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003796_s10846-021-01450-x-Figure1-1.png", "caption": "Fig. 1 The novel octocopter airframe and reference frames configuration", "texts": [ " Except considering the uncertainty of the mass and moment of inertia, we also consider the situation that the thrust and drag coefficient values of rotors are unknown and vary within a certain range. And we also verify the stability of the control system in theory based on Lyapunov stability theory, and have introduced the way to improve the control performance by adjusting only one design parameter. Finally, multiple simulations are carried out to confirm the feasibility of the novel multirotor and validity of the control method. As shown in Fig. 1, the novel fully-actuated multirotor UAV model with 8 rotors is utilized, part 1 is rotor, part 2 is motor, part 3 is arm, part 4 is body frame, and part 5 is supporting feet. Moreover, two main reference frames: the earth fixed frame OI associated with unit vector basis (xI , yI , zI ) and body fixed frame OB associated with the unit vector basis (xb, yb, zb) fixed at the center of mass of the octocopter UAV, as shown in Fig. 1. The position of center mass is denoted by the vector p = [x, y, z]T, where this vector is expressed with respected to the inertial frame OI . The attitude of this octocopter is denoted by Euler angle \u0398 = [\u03c6, \u03b8, \u03c8]T, where \u03c6 denotes roll, \u03b8 denotes pitch, and \u03c8 denotes yaw. The definitions of the utilized symbols are given in Table 1. Different from the motor arrangement of standard multirotor, the motors of this novel multirotor are fixed at the end of each arm with an included angle \u03b8k inward or outward, as shown in Fig. 2. As show in Fig. 1, the rotors with number 1,3,5,7 are tilted outward, and the rotors with number 2,4,6,8 are tilted inward. It should be noted that this structure is designed to achieve 6-DOF independent motions, although this UAV is equipped with eight rotors and it is an over-actuated system, its structure is symmetrical, which is conducive to the operation of the operator, and this design allows this UAV to have more power reserve. Moreover, different from the servo actuator structures proposed in [25\u201327], it does not need servomechanism to adjust motor angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003079_0954407020947494-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003079_0954407020947494-Figure3-1.png", "caption": "Figure 3. System responses at the subcritical Hopf bifurcation point of the original system (a2 = \u20133.231) when K1 = 100: (a) time history responses of the wheel hub speed and (b) phase diagram.", "texts": [ " Then, under the action of the designed torsional vibration stabilization controller, the original system is asymptotically stable at the origin. Taking the linear control gain of the designed torsional vibration stabilization controller as K1=100, according to the solution method in the \u2018\u2018Torsional vibration dynamic modeling\u2019\u2019 section, one can obtain that the torsional vibration stability zone of the controlled system is [\u201320.031 \u20133.044], which is obviously larger than that of the original system. Figure 3(a) and (b) shows the time history responses of the wheel hub speed and the system phase diagram at the subcritical Hopf bifurcation point of the original system when K1=100, respectively. It is seen that the subcritical Hopf bifurcation point of the original system becomes a stable point, and the output is asymptotically convergent. According to equation (17), similar to the new torsional vibration stability zone of the controlled system, the supercritical Hopf bifurcation occurs at a2=\u2013 20.031 and the subcritical Hopf bifurcation occurs at a2=\u20133" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003394_s10846-021-01352-y-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003394_s10846-021-01352-y-Figure1-1.png", "caption": "Fig. 1 Fixed-wing MAV longitudinal motion", "texts": [ " In Sections 3 and 4, the Simulink simulation block diagram and the MAV\u2019s experimental results are shown, respectively. The dynamic wind tunnel tests with a pitch degree of freedom are discussed in Sections 5. System identification was presented in section 6. In Section 7, results and discussion are presented. Attitude controller was explained in section 8. Finally, the summary and conclusions are shown in Section 9. The nonlinear longitudinal dynamics simulation model is described in this section, which includes the MAV\u2019s equations of motion and the Simulink model design. Figure 1 shows the MAV\u2019s motion in the longitudinal direction. The nonlinear equations of motion of the fixed-wing MAV are presented in the longitudinal direction by the following equations [33\u201339]. u\u0307 \u00bc Fx m \u2212qw\u2212gsin\u03b8 \u00f01\u00de w\u0307 \u00bc Fz m \u00fe qu\u00fe gcos\u03b8 \u00f02\u00de q\u0307 \u00bc M Iyy \u00f03\u00de \u03b8\u0307 \u00bc q \u00f04\u00de where Fx and Fz are the forces in the x and z directions relative to the body axes, respectively; u and w are X-axis and Z-axis velocity, respectively; g is the gravity acceleration, q is the body axis pitch rate, \u03b8 is the angle between the body axis and the horizontal line, and Iyy is the mass moment of inertia of the MAV" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001620_s12206-015-0908-1-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001620_s12206-015-0908-1-Figure8-1.png", "caption": "Fig. 8. Static stress and strain state under low frequency cycle.", "texts": [], "surrounding_texts": [ "During the operation of an aero-engine, the turbine disc is subjected to a mixed load: centrifugal forces, thermal stresses, aerodynamic forces and vibratory stresses. Of course, high speed results in large centrifugal forces and high thermal gradients result in thermal stresses. Among them, the aerodynamic forces and vibratory stresses have little effect on the static strength of the turbine disc. Therefore, when analyzing the turbine disc with finite element method, the centrifugal forces and thermal stresses are the main consideration. The speed spectrum of the turbine disc is determined by the flight mission, and it consists of three parts [11]: low frequency cycle, full throttle cycle and cruise cycle. Any speed spectrum can be considered as a combination of these three basic cycles. The speed spectrum of the turbine disc is shown in Table 3. The temperature spectrum is derived based on the measurement data. In this study, the temperature spectrum of the turbine disc was loaded on the three-dimensional model by ANSYS parametric design language. For each basic cycle mentioned above, there are 100000 temperature data points of the turbine disc. Table 4 shows part of the temperature data points under full throttle cycle, where X, Y and Z represent the coordinate value of a point of the three-dimensional model." ] }, { "image_filename": "designv11_22_0002762_j.euromechsol.2020.104125-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002762_j.euromechsol.2020.104125-Figure1-1.png", "caption": "Fig. 1. (a) Ball travel over the defect (b) Typical acceleration response of defective raceway.", "texts": [ " For degradation assessment of REB, existing fault severity becomes a guiding criterion, and one of the measures of determining the fault severity is the size of the defect on bearing components. Several studies have proposed methods to compute the defect size using signals recorded over bearings or obtained theoretically. While most of the studies use a vibration signal for obtaining defect size, few studies have used an acoustic signal for the intended task. The typical vibration response of a defective bearing is shown in Fig. 1(b) and corresponding events have been shown in Fig. 1(a). Epps and McCallion (1994) reported that the events occurring during raceway-rolling element interaction are observed in the acceleration signal. They named two points as the point of entry (point B in Fig. 1) when the element enters the defect (leading edge of the defect), and point of exit (point C in Fig. 1) when the rolling element impacts the exit end of the defect (trailing edge). For ease of understanding, in Fig. 1, the rolling element is shown to make exclusive contact with the leading-edge and trailing edge separately. At the same time, most of the studies reported in the literature assume that the element makes contacts with leading and trailing edge simultaneously. When the rolling element rolls off the leading edge, the rolling element seizes to support the load, and as a result, the element gets destressed (Sawalhi and Randall, 2011). The instant when the element crosses the trailing edge, it gets re-stressed (point D). Momentary increase in acceleration is observed at the destressing instant and shown as point B in Fig. 1(b). * Corresponding author. E-mail addresses: apatil@me.iitr.ac.in (A.P. Patil), bhanu.mishra@me.iitr.ac.in (B.K. Mishra), s.harsha@me.iitr.ac.in (S.P. Harsha). Contents lists available at ScienceDirect European Journal of Mechanics / A Solids journal homepage: http://www.elsevier.com/locate/ejmsol https://doi.org/10.1016/j.euromechsol.2020.104125 Received 28 June 2020; Received in revised form 10 September 2020; Accepted 14 September 2020 European Journal of Mechanics / A Solids 85 (2021) 104125 When the rolling element impacts the trailing edge, the acceleration level rises (point C). The re-stressing event (point D in Fig. 1(a)) causes a rise in acceleration but is often masked by the transient caused by impact. Further, it was reported that the entry event is a low-frequency event, and the exit event is a high-frequency event (Sawalhi and Randall, 2011). The time between the point of entry (point B) and point of exit (point C) was initially termed as the time to impact (TTI). The initial studies on defect size determination aimed at computing defect size from TTI. For this, a three-stage signal processing method was proposed by Sawalhi and Randall (Sawalhi and Randall, 2011): the first stage is pre-whitening of signal, the second stage balances the frequency content of two events, and the third stage computes defect size from a squared envelope of the pre-processed signal", " Then the experimental study to correlate the peak acceleration and force in the load zone is presented in the middle part of section 2. The last part of section 2 describes the signal processing based methodology to locate entry and exit events along with methods to handle any complicacy arising in it. Defect and bearing specifications, along with the experimentation carried out is presented in section 3. Section 4 discusses the results of the proposed method, and the paper is concluded in section 5. The rolling element-spall interaction shown in Fig. 1(a) is reproduced in Fig. 2 for a better illustration of the interaction. In this study, it is assumed that the rolling element does not interact with the bottom surface of the defect. The interaction of defect and rolling element, as shown in Fig. 2 (a) and (b), can be understood as the following series of events:(p, P) normal contact between the rolling element and the raceways, (q, Q) is the start of the interaction between the leading edge of the defect and the rolling element (Fig. 2(e)), (r, R) initiation of rotation of the rolling element about the leading edge, (s, S) complete destressing of the rolling element, and, (t, T) impact between trailing edge of the defect and the rolling element. Rolling element central positions A, B, and C from Fig. 1 are shown as q, s, and t respectively in Fig. 2 (a) and A.P. Patil et al. European Journal of Mechanics / A Solids 85 (2021) 104125 (b). Time spacing between rolling element center positions are defined (Luo et al., 2019a) as the time from entry to complete destressingtedand Ted, time from start to complete destressing tsd and Tsd, time from complete destressing to impact tdi and Tdi, where tis for outer race defect and is T for outer race defect. In the work of Luo et al. (2019a), the time separation tsd(or Tsd) was found out analytically using a physics-based approach, while tdi (or Tdi) was found out from recorded signal", " 7(j), (k), and Fig.7(l) respectively for 1200 rpm shaft speed. It can be concluded from Fig. 7 that the proposed method of obtaining force from peak acceleration values works satisfactorily at different speeds and defect sizes. This force is then used to analytically estimate the angles \u03c8d and \u03b1dwhich are used in defect size estimation. This further affirms that the analogy of estimating force from peak acceleration can be used in the proposed method of spall size estimation. For the purpose of generalization, Fig. 1(a) is referred to while extracting the point of complete destressing (center at B) and point of impact (center at C). As discussed at the end of section 2.1, the time A.P. Patil et al. European Journal of Mechanics / A Solids 85 (2021) 104125 duration between points B and C is to be determined directly from the measured signal. This section explains the methodology to locate the points B and C on the signal, and thus determine the time separation between the destressing to impact events. These events are to be located for each interaction between the rolling element and the defect", " If such peaks are found, the roots obtained using Ridder\u2019s method are analysed and the root closest to point B and any of the peaks with acceleration value more than 0.6amax is assumed to be point C. In Fig. 10(e) three roots square marked as 1, 2, and 3 are near to acceleration peaks having acceleration more than 0.6amax (or Tp) are located using Ridder\u2019s method. But, the root nearest to point B is 1; hence point C is chosen as a point with time corresponding to root 1. While computing the defect size, the pattern of the signal when the rolling element passes over the defect, is assumed to be similar to what is shown in Fig. 1(b), showing entry at point B and exit at point C. But, when the rolling element negotiates inner race defect, two types of patterns are observed in the signal because the position of the vibration measuring sensor is fixed, and the position of defect keeps on changing during the rotation of the inner race. Due to this, some impulses are similar to those shown in Fig. 1(b), and some are observed in inverted form. One such instance with a normal pattern of SOI is shown in Fig. 11 (b), while the inverted pattern of SOI is shown in Fig. 11(d). The corresponding position of the defect and sensor location is shown in Fig. 11 (a) and (c), respectively. When it is aimed to determine the defect size automatically, the change in the pattern of the signal may lead to wrong results. Hence a methodology is devised to detect the pattern and reverse the amplitudes if it is as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001689_0040517514547210-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001689_0040517514547210-Figure1-1.png", "caption": "Figure 1. A symmetric geometric model of spinning triangles in Siro-spinning.", "texts": [ " Three different cases where the initial strains of the shortest boundary fiber, the central fiber, and the virtual fiber on the direction of yarn load are set as zero are discussed. Then, taking the spinning triangle in Ne40 (14.6 tex) cotton Sirospun yarn as an example, the fiber tension distribution and fiber torsion distribution in the corresponding Siro-spinning triangle with and without fiber buckling are numerical simulated. A symmetric geometric model of spinning triangles in Siro-spinning is shown in Figure 1. It is easy to see that there are three spinning triangles, including two primary spinning triangles and one final spinning triangle, in Siro-spinning. Corresponding geometric models of the left and right primary spinning triangles and one final spinning triangle are shown in Figures 2\u20134, respectively. Here, w and h are the width and height of the primary triangle, is the apex angle of the primary triangle, /2 is the inclination angle of the yarn spinning tension, i is the angle between the ith fiber and the central fiber, and n is the fiber number in each fiber strand, that is, there are 2n fibers in the Siro-spinning triangle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003489_ieeeconf49454.2021.9382622-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003489_ieeeconf49454.2021.9382622-Figure3-1.png", "caption": "Fig. 3. Throwing Food Unit", "texts": [ " The mechanism is made smaller and lighter by hooking a gripping rod on the edge of the cooking vessel. A conceptual diagram of this mechanism is shown in Fig. 2 . 2) Throwing Food Unit In order to throw the ingredients into the cooking container at the right time, it is necessary to place each ingredient separately. This unit has a servo motor mounted on four food containers and a mechanism that allows the food to be fed into a slider-mounted feeder to enter the cooking containers. This mechanism is shown in Fig. 3. 3) Microwave Oven Unit In several cases, the cooking process involves heating operation. Microwaves are used in this prototype be cause they can evenly heat the material. This unit mod ifies the door of a typical microwave oven by mounting a slider onto it. This allows the robot to control the temperature and heating time of the material. Although the above units can be replaced in principle, in this study because of the nature of the mechanism of the throwing food unit, we defined the microwave oven unit as the first layer, which is attached to the upper part of the transport unit" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure1-1.png", "caption": "Fig. 1. Initial geometry of node used for BESO design (dimensions are in millimetres).", "texts": [ " Other elements which are excluded from BESO process form non-design domain. In this paper, the BESO code for optimising the structural node is linked to ABAQUS as a structural analysis engine. BESO parameters, including evolutionary ratio, volume fraction and filter radius, are set to 2%, 95%, and 3 mm respectively. Laplacian smoothing algorithm is then applied to the optimised node to smooth rough surface caused by element removal. The geometry of the node, which is used as the initial model in BESO design, as well as its dimensions are shown in Fig. 1. In this model, the non-design domain consists of small rings around the bolt hole. Two symmetrical load cases applied in the design process include axial compressive loads and out-of-plane bending moment. Fig. 2 shows the rendered perspective views of the nodes designed for symmetrical bending moment and axial compressive forces. By comparing the BESO designs for axial forces (axial node) and out-of-plane bending moment (bending node), it can be clearly observed that the optimised planar topologies for both nodes are almost identical for top and bottom planes" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003851_tsmc.2021.3103838-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003851_tsmc.2021.3103838-Figure9-1.png", "caption": "Fig. 9. Trajectory of (35) with initial condition x(0) = (0.5, 0.2, 0.3)T .", "texts": [], "surrounding_texts": [ "For the event-triggered control schemes, it is often a challenging issue to check Zeno behaviors, which are not desired to occur. Based on the definition in [47], it suffices to derive a positive low bound of the interexecution times to exclude the Zeno behavior. Thus, in the following, the positive lower bound of the interexecution {tik+1 \u2212 tik} for the interconnected system (1) under the asynchronous event-triggered control strategy will be provided in Theorem 2. Theorem 2: Consider the interconnected system with discontinuous interactions (1), for any initial conditions, if all conditions in Theorem 1 hold, the interexecution time {tik+1 \u2212 tik} is lower bounded by a positive constant T\u2020 as follows: T\u2020 = min i=1,2,...,N { 1 \u03b7i ln ( \u03b5i\u03b7i ( bi \u2212 NC ij ) \u03b6i + 1 )} (25) where \u03b7i = ki + (\u03b1im/n)2[(m\u2212n)/2n]V(0)[(m\u2212n)/2n] + (\u03b2ip/q)2[(p\u2212q)/2q]V(0)[(p\u2212q)/2q], \u03b6i = 2(1/2)\u03b7i(ki \u2212 a\u0303i \u2212 L ij \u2212 h\u0303i)V(0)(1/2) + \u03b7i(bi \u2212 NC ij)+ \u03b7i\u03b1i2(m/2n)V(0)(m/2n) + Authorized licensed use limited to: MEHRAN UNIV OF ENGINEERING AND TECHNOLOGY. Downloaded on September 03,2021 at 04:19:23 UTC from IEEE Xplore. Restrictions apply. \u03b7i\u03b2i2(p/2q)V(0)(p/2q)). Then, it can be included that the Zeno behaviors will not be exist. Proof: According to (8) and (9), it is clear that ei(t) = \u03c7i(t)\u2212 ui(t) = kixi ( tik )+ bi\u03bdi ( tik )+ \u03b1i\u03bdi ( tik )\u2223\u2223xi ( tik )\u2223\u2223m n + \u03b2i\u03bdi(t) \u2223\u2223xi ( tik )\u2223\u2223 p q \u2212 kixi(t)\u2212 bi\u03bdi(t) \u2212 \u03b1i\u03bdi(t)|xi(t)|m n \u2212 \u03b2i\u03bdi(t)|xi(t)| p q . (26) Let \u03c8i(t) = \u2016ei(t)\u2016. Then, considering the derivative of \u03c8i(t) for [tik, tik+1), we can obtain that \u03c8\u0307i(t) = d\u2016ei(t)\u2016 dt \u2264 ( ki + \u03b1im n \u2016xi(t)\u2016m n \u22121 + \u03b2ip q \u2016xi(t)\u2016 p q \u22121 ) \u2016x\u0307i(t)\u2016. (27) Since V\u0307(t) \u2264 0, according to lemma 2, the following inequality is derived: \u03c8\u0307i(t) \u2264 ( ki + \u03b1im n 2 m\u2212n 2n \u22121V(0) m\u2212n 2n + \u03b2ip q 2 p\u2212q 2q V(0) p\u2212q 2q ) \u2016x\u0307i(t)\u2016 = \u03b7i ( a\u0303i\u2016xi(t)\u2016 + L ij\u2016xi(t)\u2016 + C ij + h\u0303i\u2016xi(t)\u2016 + ui(t) ) . According to (9), the control inputs can be rewritten as ui(t) = \u2212ei(t)+ \u03c7i(t). Then, we can obtain that \u03c8\u0307i(t) \u2264 \u03b7i ( \u2016ei(t)\u2016 + ( ki \u2212 a\u0303i \u2212 L ij \u2212 h\u0303i ) \u2016xi(t)\u2016 + ( bi \u2212 NC ij )+ \u03b1i\u2016xi(t)\u2016m n + \u03b2i\u2016xi(t)\u2016 p q ) \u2264 \u03b7i\u2016ei(t)\u2016 + 2 1 2 \u03b7i ( ki \u2212 a\u0303i \u2212 L ij \u2212 h\u0303i ) V 1 2 (0) + \u03b7i ( bi \u2212 NC ij )+ \u03b7i\u03b1i2 m 2n V m 2n (0)+ \u03b7i\u03b2i2 p 2q V p 2q (0) ) . (28) From the definition of \u03b6i, (28) can be rewritten as \u03c8\u0307i(t) \u2264 \u03b7i\u2016ei(t)\u2016 + \u03b6i. (29) Then \u03c8i ( tik+1 ) = \u2225\u2225ei ( tik+1 )\u2225\u2225 \u2264 \u03b6i \u03b7i e \u03b7i ( tik+1\u2212tik ) \u2212 \u03b6i \u03b7i . (30) On the other hand, for subsystem Si, when the condition in (10) is violated, the event is triggered to update the controller. Obviously \u2225\u2225ei ( tik+1 )\u2225\u2225 \u2265 \u03b5i ( ki \u2212 a\u0303i \u2212 L ijN \u2212 h\u0303i ) \u2016xi(t)\u2016 + \u03b5i ( bi \u2212 NC ij ) . (31) Together with (30) and (31), we can obtain that ( ki \u2212 a\u0303i \u2212 L ijN \u2212 h\u0303i ) \u2016xi(t)\u2016 + \u03b5i ( bi \u2212 NC ij ) \u2264 \u03b6i \u03b7i e \u03b7i ( tik+1\u2212tik ) \u2212 \u03b6i \u03b7i . (32) By straight calculation, the interexecution time {tik+1 \u2212 tik} can be derived as { tik+1 \u2212 tik } \u2265 1 \u03b7i ln ( \u03b5i\u03b7i \u03b6i ( ki \u2212 a\u0303i \u2212 L ijN \u2212 h\u0303i ) \u2016xi(t)\u2016 + \u03b5i\u03b7i ( bi \u2212 NC ij ) \u03b6i + 1 ) . (33) According to the condition ki > a\u0303i\u2212L ijN\u2212h\u0303i in Theorem 1 and the positive parameters \u03b5i, \u03b7i, and \u03b6i, it can be concluded that [(\u03b5i\u03b7i)/\u03b6i] ( ki \u2212 a\u0303i \u2212 L ijN \u2212 h\u0303i )\u2016xi(t)\u2016 > 0. Thus, the following inequation can be achieved by relaxing the righthand side of (33) by a smaller positive one, i.e., ([\u03b5i\u03b7i(bi \u2212 NC ij)]/\u03b6i)+ 1 { tik+1 \u2212 tik } \u2265 T\u2020 i = 1 \u03b7i ln ( \u03b5i\u03b7i ( bi \u2212 NC ij ) \u03b6i + 1 ) > 0. (34) Therefore, it is easy to see that the interexecution time {tik+1 \u2212 tik} determined by the event-triggered scheme (10) is lower bounded by a positive constant T\u2020 shown as (25). The proof is completed. Remark 7: For the interconnected system with discontinuous interactions in this article, a novel controller (7) is proposed, which contains two main points. On the one hand, in order to deal with the residual term resulted from the discontinuous function fij(\u00b7), discontinuous term \u2212bi\u03bdi(tik) is designed. Actually, that is an important factor to deal with the discontinuous interaction fij and guarantee the fixed-time stabilization of the concerned systems (1). From many existing work (see [16], [42] and references therein), it can be derived that adding a discontinuous term is beneficial for handling differences between the Filippov solutions of discontinuous dynamical systems. On the other hand, aiming at realizing the stabilization of system (1) in a fixed time, \u2212\u03b1i\u03bdi(tik)|xi(tik)|(m/n) and \u2212\u03b2i\u03bdi(tik)|xi(tik)|(p/q) are introduced. Through the Lyapunov function approach, the inequation (24) is derived in the proof of Theorem 1. Thus, on the pioneer work of Polyakov [15], it can be seen that the system states will enter into the circle region with radius one within a fixed time, and after the system states enter into the region. Consequently, they will achieve exact convergence to the origin with a finite time, which is independent to the initial conditions. Remark 8: It is obvious that the event-triggered scheme (10) reflects the close relationship between triggering times and the value of threshold parameter \u03b5i. That is, the larger \u03b5i is chosen, the less events will be triggered and fewer numbers of sample data packets need to be transmitted. Consequently, in practical applications, when the control gains and related parameters are selected properly, the controller can be updated at a lower frequency. In addition, the proof of Theorem 2 has verified that the lower of interexecution times is bounded by a positive constant (25). This guarantees the control executions to be achieved only after a finite-time interval instead of infinite one, i.e., Zeno behavior [47]. Remark 9: In Theorem 1, stability criterion has been developed to ensure the fixed-time stabilization of interconnected system (1). In order to explain the solution steps of controller (7), we give a feasible design procedure of the related parameters, such as ki, bi, \u03b1i, \u03b2i, etc. IV. ILLUSTRATE EXAMPLE In order to verify the effectiveness of the control strategy and results obtained in this article, two examples are introduced to in this section. Authorized licensed use limited to: MEHRAN UNIV OF ENGINEERING AND TECHNOLOGY. Downloaded on September 03,2021 at 04:19:23 UTC from IEEE Xplore. Restrictions apply. Algorithm 1 Design of Feedback Controller (7) 1: Initialize the system parameters, including Ai,N, ni. 2: According to the interactions fij(xj(t)), choose lij, and cij, which satisfy Assumption 2. According to the external disturbance gi(xi(t)), choose \u03c2i and Hi, which satisfy \u2016gi(xi(t))\u2016 \u2264 \u03c22 i \u2016Hixi(t)\u2016. 3: Choose the control gains \u03b1i, \u03b2i and positive odd integers m, n, p, q, which satisfy \u03b1i > 0, \u03b2i > 0, m > n, p < q. 4: Choose the control gains ki and bi. 5: Check the conditions ki > a\u0303i \u2212NL ij \u2212 h\u0303i, and bi > NC ij in Theorem 1 hold or not. If yes, terminate the program and output the solutions of \u03b1i, \u03b2i, m, n, p, q, ki, bi; otherwise, go back to Step 3. Example 1: Consider an interconnected system, which includes two subsystems. The system parameters are A1 = \u23a1 \u23a3 \u22121.18 1.18 0.5 0 \u2212 0.7 0 0 0 \u2212 0.168 \u23a4 \u23a6, A2 = \u23a1 \u23a3 \u22122.57 4.5 0 1 \u2212 1 1 0 \u2212 6 0 \u23a4 \u23a6. The discontinuous interconnection fij(xj(t)) = [fij1(xj1(t)), fij2(xj2(t)), . . . , fijni(xjni(t))] T is chosen as fijs ( xjs(t) ) = { 0.5 tanh ( xjs(t) )\u2212 0.01, xjs(t) > 0 0.5 tanh ( xjs(t) )+ 0.01, xjs(t) \u2264 0 where s = 1, 2, . . . , ni. According to Fig. 1, we can see that by designing the coefficients lij and cij, a discontinuous function z = f (x) should be under the linear function z\u2217. In addition, lij and cij should satisfy the restrictions in Theorems 1. According to these ingredients, the related parameters are chosen as l12 = l21 = 0.5 and c12 = c21 = 0.01, which satisfy Assumptions 1 and 2 in this article. Without considering the controller, Fig. 3 describes the state responses of subsystem S1 with initial values x1(0) = [\u22123.5 1 \u2212 1]T , and the state responses of subsystem S2 with initial values x2(0) = [2.5 \u2212 3 \u2212 1]T . From Fig. 3, we can see that all state trajectories of the considered system are not stable. In the following, for interconnected system (1) with an asynchronous event-triggered mechanism, the proposed controllers (7) are introduced to stabilize the concerned system in fixed time. According to Theorem 1, the related parameters of u1 for subsystem S1 are selected as k1 = 6.0, b1 = 0.023, \u03b11 = 0.7, \u03b21 = 0.6, \u03b51 = 0.8, 1 = 0.02, \u03c21 = 1, H1 = diag{2, 2}, m = 7, n = 3, p = 1, and q = 3. Similarly, the related parameters of u2 for subsystem S2 are selected as k2 = 10.5, b2 = 0.022, \u03b12 = 0.4, \u03b22 = 0.3, \u03b52 = 0.9, 2 = 0.01, \u03c22 = 1, and H2 = diag{2, 2}. Based on Lemma 1, it is clear that \u03b1\u0302 = \u03b1in \u2212(m/n) i N\u2212(m/n)2[(m+n)/2n], \u03b2\u0302 = \u03b2in \u2212(p/q) i 2[(p+q)/2q], m\u0302 = m + n. n\u0302 = 2n, q\u0302 = 2q, p\u0302 = p + q. According to (10) in Theorem 1, we can get the following asynchronous event-triggered conditions. For simplicity, we take i = \u03b5i(ki \u2212 a\u0303i \u2212 L ijN \u2212 h\u0303i)\u2016xi(t)\u2016 + \u03b5i(bi \u2212 NC ij). 1) For subsystem S1, \u2016e1(t)\u2016 \u2264 1, where 1 = 0.8 \u00d7 (6.0 \u2212 1.18 \u2212 2 \u2212 0.04)\u2016x1(t)\u2016+ 0.8 \u00d7 (0.023 \u2212 0.02) = 2.78\u2016x1(t)\u2016 + 0.0024 for [t1k , t1k+1). 2) For subsystem S2, \u2016e2(t)\u2016 \u2264 2, where 2 = 0.9 \u00d7 (10.5 \u2212 6 \u2212 2 \u2212 0.02)\u2016x2(t)\u2016 + 0.9 \u00d7 (0.022 \u2212 0.02) = 2.48\u2016x2(t)\u2016 + 0.0018 for [t2k , t2k+1). According to related parameters and (11) in Theorem 1, the fixed-time for two subsystems S1 and S2 can be derived as T\u2217 1 = 8.0458 (s) and T\u2217 2 = 3.4200 (s), respectively. Authorized licensed use limited to: MEHRAN UNIV OF ENGINEERING AND TECHNOLOGY. Downloaded on September 03,2021 at 04:19:23 UTC from IEEE Xplore. Restrictions apply. Then, under the same initial conditions with Fig. 3, the state responses for the interconnected system with controllers are shown in Fig. 4. From Fig. 4(a) and (b), it is clear that subsystem S1 and S2 achieve stabilization under the proposed control scheme at t = 0.5 (s) with t < T\u2217 2 . Therefore, it can be concluded that under the asynchronous event-triggered control strategy, the event-triggered conditions in 1) and 2), the interconnected system with discontinuous interactions achieves stability in a fixed time via the newly designed controller (7). Moreover, the relationship between measurement errors \u2016ei(t)\u2016 and i for two subsystems is shown in Fig. 5. It is demonstrated that the event is triggered when \u2016ei(t)\u2016 > i. The conspicuous difference between the asynchronous eventtriggered control strategy in this article and existing ones is that the event times tik among different subsystems Si are different. That is to say, for any one of subsystems, each sensor checks the states information locally, and then decide whether to trigger the transmission of sampled data to controller or not. Similarly, all the data are transmitted asynchronously over communication channels between subsystems. This asynchronous event-triggered scheme is also illustrated by the control inputs in Fig. 6(a) and (b). From Fig. 6, we can see that the control inputs for two subsystems S1 and S2 are updated at different times. On the other hand, the simulation results show the trigger times i for two subsystems are 25 and 38, which take 10% and 15.2% of the whole data, respectively. Noteworthy, \u201cthe whole data\u201d refer to the whole sampled signals that need to be sent out to the controller in the case of no any event-triggered scheme. Consequently, the asynchronous event-triggered control strategy and Theorem 1 proposed in this article are not merely to make the interconnected system stable in a fixed time, but to reduce the number of sampled-data transmissions or controllers updates. Furthermore, for subsystems S1 and S2, Table I shows the relationship between several parameters, including eventtrigger threshold parameters \u03b5i, triggered times i, average release period \u03c4i, and the percentage of data transmissions \u03bei under the asynchronous event-triggered scheme (10). \u201cAverage release period \u03c4i\u201d refers to the average value of all interexecution times {tik+1 \u2212 tik}. \u201cPercentage of data transmissions \u03bei\u201d refers to the proportion of the sampled signals under the event-triggered control scheme (10) to the one without any event-triggered scheme. Taking subsystem S1 as an example, the results of tests (a \u223c g) show that when the event-trigger threshold parameter \u03b51 is set to be 0.1, 0.4, and 0.8, the sampled data that need to be transmitted are only 25.6%, 12.4%, and 10.0%, respectively. Meanwhile, the resource Authorized licensed use limited to: MEHRAN UNIV OF ENGINEERING AND TECHNOLOGY. Downloaded on September 03,2021 at 04:19:23 UTC from IEEE Xplore. Restrictions apply. utilization has been reduced by 74.6%, 87.6%, and 90.0% via the asynchronous event-triggered control strategy, respectively. Consequently, it can be concluded that the larger event-trigger threshold parameters \u03b5i, the less sampled signals \u03bei need to be sent, the larger average release period \u03c4i. The frequency of control update and the resources economized by the asynchronous event-triggered control strategy have an inverse proportional relationship with the event-trigger threshold parameter \u03b5i. In order to verify the superiority of our control scheme, a graphic illustration Fig. 7 is also given. Example 2: In this example, the fixed-time synchronization issue of interconnected Chua\u2019s circuit system is investigated, which is interacted by five single ones. The master system, i.e., the isolated Chua\u2019s circuit, is described as s\u0307(t) = As (t)+ Bg(s(t))) (35) where s(t) = (s1(t), s2(t), s3(t))T , B = 54/7diag{1, 0, 0}, g(s(t)) = 0.5(|s1(t)+ 1| \u2212 |s1(t)\u2212 1|, 0, 0)T A = \u23a1 \u23a3 \u2212 19 7 9 0 1 \u2212 1 1 0 \u2212 14.28 0 \u23a4 \u23a6. Under the initial condition s(0) = (0.5, 0.2, 0.3)T , the structure and trajectory of (35) are shown in Figs. 8 and 9. The overall interconnected system, i.e., the slave system, is described as x\u0307i(t) = Axi(t)+ Bgi(xi(t))+ 5\u2211 j=1 \u03c8i ( xj(t) )+ ui(t) ) (36) where i = 1, 2, 3, 4, 5, xi(t) = (xi1(t), xi2(t), xi3(t))T , \u03c8i(l) = 0.35 tanh(l) \u2212 0.02sign(l) represents the interactions between the ith sub-Chua\u2019s circuit and the other four ones. As we all know, the fixed-time synchronization issue between (35) and (36) can be regarded as the fixed-time stabilization problem of an error system, which is described as ri(t) = xi(t) \u2212 s(t). Thus, the suitable controller (7) and event-triggered scheme (10) will be designed to achieve the stabilization of system ri(t) in the following. On basis of some conditions in Theorem 1, the related parameters are chosen as k1 = 15.6, k2 = 15.4, k3 = 14.9, k4 = 15.6, k5 = 15.4, m = 7, n = 3, p = 1, q = 3, \u03b5i = 0.6, bi = 0.03, \u03b1i = 0.4, \u03b2i = 0.3, and \u03b7i = 5 for i = 1, 2, . . . , 5. Thus, under the initial conditions x1(0) = (\u22126,\u22123, 5)T , x2(0) = (\u22124.8, 6.2,\u22127.7)T , x3(0) = (7.8,\u22126.7, 3.5)T , x4(0) = (\u22126.5, 7.0,\u22123.2)T , and x5(0) = (6.4,\u22127.2, 3.1)T , the trajectory of the error system between (35) and (36) is shown in Fig. 10. From Fig. 10, it can be seen that the stabilization of the error system is achieved at t = 0.5, which is smaller than the estimate of the settling time T = 7.245 by Theorem 1. Thus, it can be concluded that under the designed controller and event-triggered scheme, the fixed-time synchronization of Chua\u2019s circuit system can be realized. By defining i = \u03b5i ( ki \u2212 a\u0303i \u2212L ijN )\u2016ri(t)\u2016+ \u03b5i(bi \u2212 NC ij), the fluttering of the error part \u2016e(t)\u2016 and the threshold part i is shown in Fig. 11. Fig. 11 indicates that the time sequences and necessary events are generated when the error part surpasses the threshold part. That is to say, unnecessary date releases, transmissions, and control updates can be avoided in stabilizing the considered system. Consequently, the obtained results are verified to be beneficial to further reduce the amount of samplings or events, and then the limited bandwidth and recourses can be saved. Finally, the interexecution times under the designed event-triggered scheme are drawn in Fig. 12. Obviously, we can easily find the maximum interexecution time 0.076 s and the minimum one 0.003 s, which guarantees that Zeno behaviors are excluded in this example. Authorized licensed use limited to: MEHRAN UNIV OF ENGINEERING AND TECHNOLOGY. Downloaded on September 03,2021 at 04:19:23 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure5.17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure5.17-1.png", "caption": "Figure 5.17 Segmental cylindrical gauge", "texts": [ " A drift slot or hole is provided near one end of the handle to enable the gauging members to be removed when replacement is necessary. The GO and NOT GO gauges may be in the form of separate \u2018single-ended \u2019 gauges or may be combined on one handle to form a \u2018double-ended \u2019 gauge, Fig. 5.16. Figure 5.16 Plug and ring gauges Since the GO gauging member must enter the hole being checked, it is made longer than the NOT GO gauging member, which of course should never enter. Large gauges which are heavy and difficult to handle do not have a full diameter but are cut-away and are known as segmental cylindrical gauges, Fig. 5.17. Adjustable gap gauges, Fig. 5.19, consist of a horseshoe frame fitted with plain anvils the spacing of which can be adjusted to any particular limits required, within the range of the gauge. Setting to within 0.005 mm of a desired size is possible with well-made adjustable gauges. When checking a shaft, the GO gap gauge should pass over the shaft, under its own weight when the axis of the shaft is horizontal or without the use of excessive force when the axis of the shaft is vertical. A cylindrical ring GO gauge should pass over the complete length of the shaft without the use of excessive force" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002475_s40430-020-02295-5-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002475_s40430-020-02295-5-Figure6-1.png", "caption": "Fig. 6 Comparison between calculated and experimental results at the speed of 18,000\u00a0r\u00a0min\u22121", "texts": [ "5\u00a0mV\u00a0Pa\u22121 were placed at the measuring points. The sound pressure signals of each measuring point were collected by a data collector (INV3018C). The sampling frequency and sampling time were 51.2\u00a0kHz and 10\u00a0s, respectively. The collected data were sent to the computer for further processing the sound pressure data and analyze the radiation sound field characteristics of the FCACBB. With the spindle spun counterclockwise at 18,000\u00a0r\u00a0min\u22121, the calculated values and experimental results in the circumferential direction are shown in Fig.\u00a06. As shown in Fig.\u00a06, it can be seen that the calculated and experimental SPLs have similar change trend along the circumferential direction, and all experimental results are relatively low compared with calculated values. The primary reason for the case is that the transmission of the radiation noises of the FCACBB is only considered under the condition of the ideal sound propagation and equivalent analysis in the calculated model, while the sound waves get more attenuation and absorption due to the barrier of housing of the motorized spindle during the experiment" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003429_j.tws.2021.107585-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003429_j.tws.2021.107585-Figure6-1.png", "caption": "Fig. 6. Edge loads for the pressure vessel with torispherical ends.", "texts": [ " \ud835\udf111 = 0 following boundary conditions are assumed: ( ) ( ) \ud835\udc51\ud835\udc63 \ud835\udf111 = 0, \ud835\udc51\u210e \ud835\udf112 = 0, \ud835\udf171(\ud835\udf111) = 0. (40) The boundary conditions provided in Eqs. (40)\u2013(42) consider separated shell structures. To achieve structural compatibility in the junctions of deformed pressure vessel additional boundary conditions have to be applied. The compatibility equations i.e. boundary conditions for the junction of the ellipsoidal and cylindrical shell (Fig. 5) have the form: \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc522 ) = \ud835\udc51\ud835\udc63 ( \ud835\udf09\ud835\udc502 ) , \ud835\udf171 ( \ud835\udf11\ud835\udc522 ) = \ud835\udf171 ( \ud835\udf09\ud835\udc502 ) . (43) For the pressure vessel with torispherical dished ends (Fig. 6) three shells constituting two junctions have to be considered, therefore: \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc602 ) = \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc611 ) , \ud835\udf171 ( \ud835\udf11\ud835\udc602 ) = \ud835\udf171 ( \ud835\udf11\ud835\udc611 ) , \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc612 ) = \ud835\udc51\ud835\udc63 ( \ud835\udf09\ud835\udc502 ) , \ud835\udf171 ( \ud835\udf11\ud835\udc612 ) = \ud835\udf171 ( \ud835\udf09\ud835\udc502 ) . (44) After applying the boundary conditions described in Eqs. (40)\u2013(42) to the displacement functions it is possible to formulate the total potential energy as a function of the remaining \ud835\udc4e\ud835\udc56 and \ud835\udc4f\ud835\udc56 parameters. Importantly, one must consider the number of parameters \ud835\udc4e\ud835\udc56 and \ud835\udc4f\ud835\udc56 in the displacement functions hereinafter denoted \ud835\udc5b\ud835\udc5d, as well as number of independent variables \ud835\udc5b in the total potential energy expression", " The unknown parameters \ud835\udc50 in the displacement functions can be then obtained for both of the shells: \ud835\udc53\ud835\udc50 = { \ud835\udc4b\ud835\udc501, \ud835\udc4c\ud835\udc501, \ud835\udc4d\ud835\udc501, 0, 0, 0, \ud835\udc5d }\ud835\udc47 , (59) \ud835\udc50\ud835\udc50 = \ud835\udc48\ud835\udc50 \u22121\ud835\udc4a\ud835\udc50\ud835\udc53\ud835\udc50 , (60) \ud835\udc53\ud835\udc52 = { 0, 0, 0, \ud835\udc4b\ud835\udc522, \ud835\udc4c\ud835\udc522, \ud835\udc4d\ud835\udc522, \ud835\udc5d }\ud835\udc47 , (61) \ud835\udc50\ud835\udc52 = \ud835\udc48\ud835\udc52 \u22121\ud835\udc4a\ud835\udc52\ud835\udc53\ud835\udc52. (62) Torispherical dished end, unlike ellipsoidal, is characterized by the discontinuous radii of curvature along its meridian. Such fact implies that spherical and toroidal shells have to be considered separately. The edge loads for the pressure vessel with torispherical ends are presented in Fig. 6. The compatibility equations must include both junctions simultaneously, namely spherical to toroidal and toroidal to cylindrical. Such approach implies the both edges of the toroidal shell are interacting with each other i.e. loads at the first edge cause displacement and rotation on the second edge and vice versa. Therefore, in the following equations, displacements and rotations on both of the edges of the toroidal shell contain all of the forces, namely \ud835\udc4b1\ud835\udc61, \ud835\udc4c1\ud835\udc61, \ud835\udc4d1\ud835\udc61, \ud835\udc4b2\ud835\udc61, \ud835\udc4c2\ud835\udc61, \ud835\udc4d2\ud835\udc61: \ud835\udeff\ud835\udc501( \ud835\udc4b1)\ud835\udc4b\ud835\udc501 + \ud835\udeff\ud835\udc501( \ud835\udc4c 1)\ud835\udc4c\ud835\udc501 + \ud835\udeff\ud835\udc501( \ud835\udc4d1)\ud835\udc4d\ud835\udc501 + \ud835\udeff\ud835\udc501( \ud835\udc5d)\ud835\udc5d = \ud835\udeff\ud835\udc612( \ud835\udc4b2)\ud835\udc4b\ud835\udc612 + \ud835\udeff\ud835\udc612( \ud835\udc4c 2)\ud835\udc4c\ud835\udc612 + \ud835\udeff\ud835\udc612( \ud835\udc4d2)\ud835\udc4d\ud835\udc612 + \ud835\udeff\ud835\udc612( \ud835\udc5d)\ud835\udc5d + \ud835\udeff\ud835\udc612( \ud835\udc4b1)\ud835\udc4b\ud835\udc611 +\ud835\udeff\ud835\udc612(\ud835\udc4c 1)\ud835\udc4c\ud835\udc611 + \ud835\udeff\ud835\udc612( \ud835\udc4d1)\ud835\udc4d\ud835\udc611, \ud835\udeff\ud835\udc611( \ud835\udc4b1)\ud835\udc4b\ud835\udc611 + \ud835\udeff\ud835\udc611( \ud835\udc4c 1)\ud835\udc4c\ud835\udc611 + \ud835\udeff\ud835\udc611( \ud835\udc4d1)\ud835\udc4d\ud835\udc611 + \ud835\udeff\ud835\udc611( \ud835\udc5d)\ud835\udc5d + \ud835\udeff\ud835\udc611( \ud835\udc4b2)\ud835\udc4b\ud835\udc612 +\ud835\udeff\ud835\udc611(\ud835\udc4c 2)\ud835\udc4c\ud835\udc612 + \ud835\udeff\ud835\udc611( \ud835\udc4d2)\ud835\udc4d\ud835\udc612 = \ud835\udeff\ud835\udc602( \ud835\udc4b2)\ud835\udc4b\ud835\udc602 + \ud835\udeff2\ud835\udc60( \ud835\udc4c 2)\ud835\udc4c\ud835\udc602 + \ud835\udeff\ud835\udc602( \ud835\udc4d2)\ud835\udc4d2\ud835\udc60 + \ud835\udeff\ud835\udc602( \ud835\udc5d)\ud835\udc5d, \ud835\udf17\ud835\udc501( \ud835\udc4b1)\ud835\udc4b\ud835\udc501 + \ud835\udf17\ud835\udc501( \ud835\udc4c 1)\ud835\udc4c\ud835\udc501 + \ud835\udf17\ud835\udc501( \ud835\udc4d1)\ud835\udc4d\ud835\udc501 + \ud835\udf17\ud835\udc501( \ud835\udc5d)\ud835\udc5d = \ud835\udf17\ud835\udc612( \ud835\udc4b2)\ud835\udc4b\ud835\udc612 + \ud835\udf17\ud835\udc612( \ud835\udc4c 2)\ud835\udc4c\ud835\udc612 + \ud835\udf17\ud835\udc612( \ud835\udc4d2)\ud835\udc4d\ud835\udc612 + \ud835\udf17\ud835\udc612( \ud835\udc5d)\ud835\udc5d + \ud835\udf17\ud835\udc612( \ud835\udc4b1)\ud835\udc4b\ud835\udc611 +\ud835\udf17\ud835\udc612(\ud835\udc4c 1)\ud835\udc4c\ud835\udc611 + \ud835\udf17\ud835\udc612( \ud835\udc4d1)\ud835\udc4d\ud835\udc611, \ud835\udf17\ud835\udc611( \ud835\udc4b1)\ud835\udc4b\ud835\udc611 + \ud835\udf17\ud835\udc611( \ud835\udc4c 1)\ud835\udc4c\ud835\udc611 + \ud835\udf17\ud835\udc611( \ud835\udc4d1)\ud835\udc4d\ud835\udc611 + \ud835\udf17\ud835\udc611( \ud835\udc5d)\ud835\udc5d + \ud835\udf17\ud835\udc611( \ud835\udc4b2)\ud835\udc4b\ud835\udc612 +\ud835\udf17\ud835\udc611(\ud835\udc4c 2)\ud835\udc4c\ud835\udc612 + \ud835\udf17\ud835\udc611( \ud835\udc4d2)\ud835\udc4d\ud835\udc612 = \ud835\udf17\ud835\udc602( \ud835\udc4b2)\ud835\udc4b\ud835\udc602 + \ud835\udf17\ud835\udc602( \ud835\udc4c 2)\ud835\udc4c\ud835\udc602 + \ud835\udf17\ud835\udc602( \ud835\udc4d2)\ud835\udc4d\ud835\udc602 + \ud835\udf17\ud835\udc602( \ud835\udc5d)\ud835\udc5d" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001790_s1064230715040139-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001790_s1064230715040139-Figure4-1.png", "caption": "Fig. 4. Terminal curves for various values of accelerating force under fixed .0R", "texts": [ " Taking into account the self similarity property, we conclude that if there is an arbitrary terminal point and there is a trajectory, which is derived as a solution of system (3.26), 0R 0 0 05R = . 1 3(0) (0) 0q q= = \u03a8 0R 0R 522 JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 54 No. 4 2015 VONDRUKHOV, GOLUBEV which connects this point to the predetermined initial point, then this trajectory is unique for this pair of points. To estimate the effect of the accelerating force on the shape of terminal curves, the terminal curves for various values of parameter c (Fig. 4) have been plotted. When parameter c decreases, we can see that the upper part of the terminal curve approaches the horizontal axis Ax. Under c = 0 we obtain a piece of the cycloid as a solution of the classical brachistochrone problem. 5. EFFECT OF VISCOUS FRICTION To derive the optimal trajectories under the action of viscous friction and a constant accelerating force, we use system (3.24). In accordance with [5], to avoid the singularity at the initial time, when the velocity is zero, we set a quasi constant accelerating force using the following coefficient: (5" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002712_dcoss49796.2020.00068-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002712_dcoss49796.2020.00068-Figure3-1.png", "caption": "Fig. 3. Different cases of line loss.", "texts": [ " From the above, we can safely infer that the PID-only controller, when it comes to controlling a robot towards autonomous tracking and following a line/path, has profound limitations due to the extreme changes of the path curvature. An extreme change in the path, such as an acute or right angle, can lead to large deviations from the set-point and consequently large errors that can\u2019t be controlled adequately, resulting in sudden loss of the line under the sensor and uncontrollable behaviour. Consider the case presented in Figure 3 where the classical PID controller leads to line loss due to extreme changes in the path curvature. One can visualize how the PID-only controller, due to it\u2019s feedback dependency, can suffer from accidental line loss. Moreover, PID also suffers from overshot and parameters setting in tracking performance [25], characteristics that create the need for a new control system, more suitable and more robust for line following and tracking. Our proposed method strives to correct these limitations by adding an anticipatory control, via the addition of an input from a new feed-forward system that can greatly enhance the ability of the PID controller to react timely in fast non-linear deviations from the set-point" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002736_jsen.2020.3022421-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002736_jsen.2020.3022421-Figure1-1.png", "caption": "Fig. 1. (a) Model of a soft hand; (b) The structure of soft finger", "texts": [ " Secondly, we analyzed the relationship between the wavelength and curvature of the fiber grating sensor using optical fiber sensing theory and shape modeling method. Finally, the deformation and shape are analyzed and simulated by means of Cosserat theory and ABAQUS software. The calculated results were compared with experimental results to verify the effectiveness of the prediction ability of soft finger\u2019s shape through FBG sensors. In order to study the grasping behavior of the soft hand, we designed a soft hand with three fingers as shown in Fig. 1(a). Fig. 1(b) shows the detailed structure of the finger. The finger skeleton is composed of four thermal-induced wires (SMA-1) to change the stiffness of the finger through the electric current for the increasing grasping force, and one super-elastic wire (SMA-2) to achieve the purpose of recovering after the bending deformation of the finger, as well as a driving wire (SMA-3), which can contract to get bending torque of the finger provided by heating electric current. The FBG is arranged on the same side of the SMA-2 wire to sense the curved shape of the finger" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000188_s11071-019-04947-1-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000188_s11071-019-04947-1-Figure7-1.png", "caption": "Fig. 7 Presentation of some kinematic elements of the collisions between two balls rolling over the surface of a sphere: a and b velocities of the contact point between two balls different radiuses; c and d impact angular velocities of the rolling balls before collision; e and f directions of the traces of rolling of two balls and impact angular velocities before and outgoing angular velocities after collision (for details and explanation, see Refs. [12\u201314])", "texts": [ " And after collision, the obtained outgoing angular velocities and angle coordinates of two contact points of balls with sphere in configuration of first collision we take as initial conditions for balls rolling over the sphere using derived nonlinear differential equation of moving, and equation of phase trajectory, up to the next position of second collision between rolling balls. And then according to the same previous procedure, we can continue investigation of the possible occurrence of the next collision between balls. This is practically a foundation of the methodology for analysis and investigation of the dynamics of vibro- impact system with two rolling balls that roll over the sphere surface. Attached Fig. 7 shows kinematics of the collision configurations of two equal balls (a and b; c and d), as well as two rolling balls with different radiuses in the collision configuration with the corresponding rolling traces (e and f), and rolling trace along parallel in circular direction, for illustrating the kinematics of the collision of two balls in rolling (see References [1,2] and [3]). When we take the task of determining coordinates of contacts of rolling balls with the sphere surface in which they roll, in the collision configuration, it is very useful to continue investigation using the phase plane method" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001347_j.ymssp.2014.06.016-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001347_j.ymssp.2014.06.016-Figure5-1.png", "caption": "Fig. 5. Three-wheeled mobile platform.", "texts": [ " Finally, the developed mobile platform construction of the platform, and requirement for mobile platform conveyance without slips, determined the geometrical structure and the adopted solutions for propulsion and steering systems. Deployed algorithm, which tailors the architecture and fosters on-line in minimizing mobile platform displacement errors, as well as\u2014very strong dynamic nonlinearity, imposed the need for highly efficient computing controller. Thereupon it was presented in a view of the designed platform (Fig. 5). As it was mentioned formerly, due to necessity of wheels slips exclusion, the differential system, connected centrally to electrical motor (Micromotors) had to be deployed. Electric line scheme was depicted (Fig. 6). The main component of developed system was based on the cRIO NI-9076 embedded controller powered by the LabVIEW. The ruggedness of cRIO controller concept by high-end and flexible architecture allows developing deterministic application both in real time processor and FPGA. The meaningful and decisive factor in the choice of mobile platform DC motors had incorporated separate driver module NI-9505, which beside PWM generation feature includes a built-in quadrature encoder interface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure21.8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure21.8-1.png", "caption": "Figure 21.8 Third angle projection", "texts": [ " It is usually the simplest and quickest way of giving an accurate representation, and dimensioning is straightforward. It is only necessary to draw those views that are essential and it is simple to include sections and hidden detail. Orthographic projection may be first angle or third angle and the system used on a drawing should be shown by the appropriate symbol. Fig. 21.7 each view showing what would be seen by looking on the far side of an adjacent view. C B C A AB Front elevation looking on arrow A Plan view looking on arrow C End elevation looking on arrow B Figure 21.7 First angle projection Fig. 21.8 each view showing what would be seen by looking on the near side of an adjacent view. This is a method of showing, on a single view, a three-dimensional picture of an object. It is therefore easier to visualise and is useful for making rough sketches to show someone what you require. Two kinds of pictorial projection are used: isometric projection and oblique projection. In isometric projection, vertical lines are shown vertical but horizontal lines are drawn at 30\u00b0 to the horizontal on each side of the vertical" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003430_09544070211004507-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003430_09544070211004507-Figure1-1.png", "caption": "Figure 1. Physical structure of quarter HM-EV system.", "texts": [ " Although these studies considered UMP of hub motor, they did not provide a UMP calculation method with concision and precision. Therefore, it is necessary to propose a concise and accurate calculation method of UMP, and introduce adjustable force to suppress longitudinal and vertical vibration of the HM-EV simultaneously. Thus, a tendegree-of-freedom quarter HM-EV system model that takes longitudinal and vertical vibration characteristics of air suspension, hub motor and tire into consideration is proposed in this article firstly. The physical structure of quarter HM-EV system is shown in Figure 1. The accuracy of the model is verified by experimental validation. Secondly, UMP calculated by BP neural networks is proposed based on this model. Finally, an active suspension controller is designed to improve HM-EV dynamic performance. This article is arranged as follows. In section \u2018\u2018Quarter HM-EV system modeling,\u2019\u2019 a ten-degree-offreedom mathematical quarter HM-EV system model is built. The BLDC hub motor is introduced into the HM-EV system. The influence of UMP on suspension system can be analyzed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001698_icelmach.2014.6960379-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001698_icelmach.2014.6960379-Figure6-1.png", "caption": "Fig. 6. Arrangement of the windings in a 3-phase system (a) and Cartesian coordinates system (b).", "texts": [ " Due to the fact that the simulated movements of the rotor are small (especially in the case of Figure 4a), an important aspect of the creation of the FEM model is to create a respectively dense finite element mesh particularly in the air gap. Examples of calculated phase voltage waveforms generated in machine windings during vibrations of rotor are shown in Fig. 5. In this case, vibrations were sinusoidal. In order to facilitate the analysis, calculated and measured values of generated voltages uA, uB, uC can be presented in Cartesian coordinates u\u03b1, u\u03b2 as shown in Fig.6. Transformation between three-phase system and the Cartesian can be done using (1)-(4) [4], graphical interpretation of space vector of voltage u and its components u\u03b1, u\u03b2 are ilustrated in Fig. 7. [ ])()()(1 3 2 2 tuatuatuu CBA ++= (1) 2 3 2 13 2 jea j +\u2212== \u03c0 (2) ( ) A uCuBu Auuu = + \u2212== \u239f \u239f \u23a0 \u239e \u239c \u239c \u239d \u239b 23 2 Re \u03b1 (3) ( ) 3 Im CuBu uu \u2212 == \u03b2 (4) 1493 As described previously, one of the investigated variants of vibration was harmonic motion of rotor (Fig. 4a). In this case the position of rotor in computational model has been described as a sinusoidal function" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003001_j.engfailanal.2020.105195-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003001_j.engfailanal.2020.105195-Figure8-1.png", "caption": "Fig. 8. Sampling from the top tube of the bicycle frame: a) general view, b) mini-specimen from the joint material, units in mm.", "texts": [ " For smaller cross-sectional areas, other failure mechanism may be observed at the initiation and propagation of fatigue cracks than in standard specimens. Very small specimens yield higher scatter of data due to random local changes in material properties. The determination of fatigue properties (base material) was supplemented by the analysis of local properties near the fatigue crack (joint material). A heat affected zone from the welded joint indicated by the discoloration of steel on the inner side of the tube can be observed in the tested node. Fig. 8 shows the sampling diagram for two areas. The figure also shows the sampling distance measured from the welded joint to the working section of the specimen. The specimens were taken symmetrically from two sides of the bicycle frame. The material data based on mini-specimens taken from the region in the immediacy of the fatigue crack gives information on actual fatigue properties and determines the approximate effect of machining and the condition of the outer material layer. The results can be treated as laboratory test results and will be compared using similar conditions (constant amplitude loading)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003373_5.0035987-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003373_5.0035987-Figure4-1.png", "caption": "FIGURE 4. Schematic diagram of electrochemical processing", "texts": [ " The combined dimensional processing method includes laser processing of the part using a pulsed fiber laser on a precision laser complex in the first stage, after which, a finally formed artificial roughness is obtained on the workpiece surface in the finishing mode by electrochemical processing [8]. At the first stage of processing (Figure 3), part 1 and fiber laser 2 are installed on the desktop 3 of a precision laser complex with a rotary device 4. Laser 2 is located at a certain focal length L, depending on the type of fiber laser. After that, energy is supplied to the laser 2 and the work surface 5 is processed. The location of the part 1 and the laser 2 relative to each other is controlled by the rotary device 4. 030004-3 At the second stage (Fig. 4) of the processing, the working part 5 of part 1 is set relative to the cathode device 6 on the working table 3 opposite the surface 5 obtained in the first stage with a fiber laser. The electrolyte 7 is fed into the space between the part 1 and the cathode device 6. Through the rectifier, current is supplied to the electrolyte, while the part 1 and the cathode device 6 are turned on according to the direct polarity scheme. When forming a surface layer subject to temperature influences, its removal occurs by electrochemical treatment, in special technological conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002725_tmag.2020.3021644-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002725_tmag.2020.3021644-Figure9-1.png", "caption": "Fig. 9. DFM-CMG prototype. (a) HSR, (b) Stationary PM ring (LSR), (c) Output rotor.", "texts": [ " In our work, to fabricate the PM ring, the ends of PM-segments are inserted into the holes in mounting seat and reinforced with glue. (ii) As discussed above, to achieve the maximum torque capability, the OM segments should be aligned with the IM teeth. While, for an assembled device, there is a misalignment between the two parts due to the installation errors. According to Fig. 7(d), the tolerance of \u03b1m-m is designed within \u00b10.5\u00b0. Moreover, for facilitation of the lamination of silicon steel sheets, the ferromagnetic segments on OM are connected with bridges in our work. The DFM-CMG prototype is shown in Fig. 9. To reduce the manufacturing cost, rectangular PMs are used. One pole of the HSR in Fig. 9(a) is composed of 4 PM-segments. Unlike the operation mode discussed in Section III, here, the PM ring (LSR) in Fig. 9(b) is fixed. And the combination of modulators in Fig. 9(c) is used as the output rotor. The 2D flux Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 20,2020 at 21:04:18 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. density distribution corresponding to the prototype is plotted in Fig. 10. It should be pointed out that the PM configuration, connecting bridges and mounting holes on OM may decrease the performances of DFM-CMG compared with the data in Table II" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001325_ssd.2014.6808802-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001325_ssd.2014.6808802-Figure4-1.png", "caption": "Fig. 4. Reference {B} and body {Q} frames", "texts": [ " Given that D is the distance from the center of the propeller to the center of gravity COG and K is a constant that relates moment and thrust of a propeller, the moments acting on the quadrotor are expressed as, Mz = (Tl + T3 - T4 - T2)K (5) (6) (7) (8) Moreover, the motors' thrusts are related to the motor's angular velocity as (9) Where c is the static thrust constant If the motors are investigated as a system, they can be studied with the voltage as inputs and angular velocity and/or torque as outputs. Fig. 3 shows the motors inputs, control signals, provided as Pulse Width Modulation (PWM) voltages and torque outputs. So far, the issue of relative frames has not been discussed. However, an aerial vehicle's orientation and position should be related to a reference frame. Fig. 4 shows the quadrotor in frame {Q} relative to a base frame {B}, Because our main objective is concerned with stabilization at a specific height, the relative position transformation and relative linear velocities will not be investigated. However, the relative angular dynamics are of interest. The rotation matrix R, which is the result of the product between three other matrices (R (<1\u00bb, R (8) and R (tV)), each of them representing the rotation of the Q frame around each one of the B axis, is formulated as [ dJcl/J R = sl/Js(Jcl/J - cepsl/J ceps(Jcl/J + sepsl/J c(Jsl/J cepcl/J + seps(J sl/J s(Jcepsl/J - sepcl/J -s(J 1 sepc(J c(Jcep (10) To find the relationship between the Euler rates [ e tjJ] and the body axis angular rates[p q r] the following equality must be satisfied, w = pt + qJ + rk = (p + e + Jj; (11) Since three consecutive revolutions makes the Euler angle sequence, meaning that the Roll angular rate takes one revolution, the Pitch angular rate takes two, and the Yaw angular rate takes three" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003480_ieeeconf49454.2021.9382668-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003480_ieeeconf49454.2021.9382668-Figure1-1.png", "caption": "Figure 1. Structure of walking assistive device.", "texts": [ "[7] also utilized the ZMP as a stability index and then adjusted the device\u2019s motion according to the ZMP location. Based on the limitation of the existing assistive devices, this paper aimed to ensure safety during the stair-ground transition when patients wearing the walking assistive devices. The method was applied to a device with spatial parallel-link mechanism firstly[8] . II. Wa l k i n g A s s is t a n c e De v ic e The proposed walking assistive device in this paper which can support the whole lower limb[7] is shown in Fig. 1. The total weight of the device is 23kg and the user does not need to bear the weight of device by gravity compensation. 4 DC motor (150W) were mounted on the hip and knee joints, and encoders were attached to measure the rotation angle of each joint. The output torque of motor is 30Nm according to the gear reduction ratio 126:1. Elderly people need to face the flat ground and stairs in daily life inevitably. But it is difficult for elderly people to walk on the stairs. A zero-moment point control method was applied to stabilize the walking motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002379_0954406220917424-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002379_0954406220917424-Figure2-1.png", "caption": "Figure 2. Schematic diagram of the tooth modification. (a) Comprehensive modification. (b) Profile modification.", "texts": [ " The normal module is 10mm, the face width is 220mm, the normal pressure angle and the helix angle are 20 and 7.493 , respectively, and the quality grade is 5. For computational convenience, the major design parameters of the example helical planetary gearing stage are listed in Table 1. Effect of tooth modification on relative displacements along line of action From a mechanics perspective, tooth modification can be regarded as an internal displacement excitation of a gear system. Therefore, the relative displacement along the LOA caused by a tooth modification is derived below. Figure 2 illustrates the schematic diagram of a helical gear tooth, in which the solid lines represent the tooth flanks without modification, and the dashed lines represent the modified tooth flanks. Herein, the tooth is modified with a profile modification and lead modification. In Figure 2, B is the face width, points O and O\u2019 are the base circle center of the involute profile and the circular arc center of the axial modification, respectively, a is the tooth profile modification amount, Ca is the axial modification amount, Ra is the crowned radius, points A and A\u2019 are the starting points of the original and the modified involutes, respectively. is the rotating angle around the Z axis at any point on the tooth surface in the XOY plane, and is the helix angle. ra is the radius of the addendum circle, rj is the radius vector of the intersection point between the original and modified involutes, rb and rb\u2019 are the base circle radii of the original and modified involutes, respectively, \u2019 is the angle between the radius vector of the starting point of the original and modified involutes, and \u2019\u2019 is the angle between the radius vector of the starting point of the original profile and the Y-axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003667_mawe.202000242-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003667_mawe.202000242-Figure1-1.png", "caption": "Figure 1. Schematic illustration of the gas metal arc welding based cladding process: (a) deposition process and (b) tensile specimen size. Arrows indicate the deposition and building direction.", "texts": [ " The cladding parameters considered for the deposition were determined by evaluating the weld bead geometry and width of the weld bead with minor delineations. The cladding parameters were welding current of 160 A, welding speed of 250 mmmin 1, welding voltage of 16.4 V, wire feed rate of 5.23 mmin 1, and deposition rate of 2.36 kgh 1. A gas mixture of 98 % Ar and 2 % CO2 was used for shielding at a steady flow rate of 20 Lmin 1. Schematic representation of the tensile specimen and gas metal arc welding based cladding process, Figure 1. After cladding, the cladded layers were face milled to obtain good surface finish and uniform thickness. Microstructure was captured for different samples according to American society for testing and materials E3-11(2017) standard along the building direction. The metallographic specimens were cleaned and polished to mirror surface finish. Consequently, they were etched electrochemically with 10 % oxalic acid solution at 10 V for 10 seconds. Tensile samples were prepared along the building direction using wire-cut electrical discharge machining process" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002102_lars-sbr.2016.43-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002102_lars-sbr.2016.43-Figure2-1.png", "caption": "Figure 2. Geometric analisys of the intersection of bubbles.", "texts": [ " / 0 1 , 2 3 / 4 0 5 6 / 0 7 , 2 3 / 4 0 6 Where , is the radius of bubble q1, 3 is the radius of bubble q2 and / 0 is the distance between the their centers, measured with the same metric used on C. For each detected intersection, the procedure adopted to avoid the creation of redundant bubbles consists on modifying the probability distribution function on the surface of the parent bubble by setting to zero the probability of sampling a new bubble center in the region of intersection. This procedure for a two-dimensional configuration space, where the bubbles are circles, is shown in Figure 2. The limit angles of the region of intersection, 8, e 83 (Fig. 2 (b)), can be easily obtained through geometric relations: 8, 8 9 8: 83 8 2 8: Where 8 is the angle between the X-axis and the vector that connects the two centers of the bubbles, and d is the distance between them: 8: ;<=+, > 3 2 ,3 9 33?@ ,@ A The points between 8, and 83 on the perimeter of the bubble 1 have their probability of being sampled imposed to zero in the probability distribution function. After withdraw this interval, a normalization must be done in order to keep the final value of the cumulative probability distribution function equal to1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003527_j.tafmec.2021.102991-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003527_j.tafmec.2021.102991-Figure1-1.png", "caption": "Fig. 1. Involute spur gear.", "texts": [ " In addition, a parametric study is carried out to determine the influence of important gear parameters on SIFs with a 3D crack at the critical region of the fillet to design the asymmetric spur gears. The asymmetric spur gear with the involute profile is developed by increasing the drive side pressure angle. Two base circles with different pressure angle are designed according to Eqs. (1) and (2). The cartesian coordinates at the radius ri of involute profile xi and yi with the center as the origin is given by (Fig. 1) [5] xi = risin\u03b8i and yi = ricos\u03b8i (1) where \u03b8i = \u239b \u239d ( \u03c0m 4 ) + \u03c0mtan\u03b1o ro + inv\u03b1o \u2212 inv\u03b1i \u239e \u23a0 (2) N.V. Namboothiri and P. Marimuthu Table 1 Parameters for convergence study. Parameter Value Module, m (mm) 1 Number of teeth, z 30 Pressure angle, \u03b1 20\u25e6 Gear ratio, i 2 Fig. 6. The variation of bending stress of symmetric gear with element size. N.V. Namboothiri and P. Marimuthu Theoretical and Applied Fracture Mechanics 114 (2021) 102991 where m \u2013 module ro \u2013 pitch circle radius ri \u2013 Radius if circle at ith position on involute profile of spur gear \u03b1o \u2013 pressure angle \u03b1i \u2013 pressure angle at the ith position of the gear The effect of variation of gear parameters on the top land thickness and contact ratio has been investigated for selecting the gear parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001483_0954406215612814-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001483_0954406215612814-Figure4-1.png", "caption": "Figure 4. Graphical application of the Euler\u2013Savary equation to both conjugate profiles (a higher-kinematic-pair) to give the instantaneously equivalent four-bar linkage O s Ol, where is the center of curvature of the trajectory of s and vice versa, and s is the center of curvature of that of , whose osculating circles are depicted by two dashed circular arcs.", "texts": [ " This feature will be clarified in the Euler\u2013Savary equation for envelopes and Aronhold theorems and the return circle sections based on the Euler\u2013Savary equation for envelopes and Aronhold\u2019s first theorem, which takes into account the properties of the return circle. The Euler\u2013Savary equation for envelopes can be obtained by referring to the sketches of Figure 3, which show the graphical determination of the centers of curvature and s, of and s, respectively, as envelopes of during the pure-rolling motions of \" on l (Figure 3a) and \" on l (Figure 3b), respectively. Thus, both conjugate profiles and s can be considered in contact at the point P of Figure 4, where the graphical construction of Figure 3 is applied again by showing that the centers of curvature and s are also, correspondingly, the centers of curvature for their relative motion represented by the centrodes l and l, whose centers of curvature are Ol and Ol , respectively. Thus, the Euler\u2013Savary equation for envelopes can be expressed as 1 P0 1 P0 l cos \u00bc 1 rl 1 rl \u00f01\u00de where is a polar coordinate of point P with respect to the ordinate axis of the canonical frame (P0, t, n), P0 and P0 l giving the positions of the centers of curvature and l with respect to P0, while rl and rl are the signed radii of curvature of l and l, respectively", " Thus, the centers of curvature and s are obtained as intersections with N , of two parallel lines passing through Ol and Ol, respectively. Aronhold theorems and the return circle The return circleR is the locus of the centers of curvature of all points at infinity of a moving plane, whose equation can be readily derived from the Euler\u2013 Savary equation 1 P0 1 P0M cos \u00bc 1 rl 1 r l \u00bc 1 \u00f02\u00de where P0M and are the polar coordinates of a generic moving point M with respect to the canonical frame (P0, t, n) shown in Figure 4, and is the corresponding center of curvature. In fact, by letting P0M!1 for all points at infinity, equation (2) leads to P0 \u00bc cos \u00f03\u00de which is the equation of the return circle R of diameter that is tangent to both centrodes l and l at the at UNIV CALIFORNIA SAN DIEGO on December 14, 2015pic.sagepub.comDownloaded from instant center of rotation P0. This circle is a mirror image of the inflection circle with respect to the tangent line t and takes the place of the inflection circle for the inverse motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001555_978-3-319-18944-4_10-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001555_978-3-319-18944-4_10-Figure1-1.png", "caption": "Fig. 1. 1-1 Leader-Follower coordinate definition", "texts": [ " Secondly, an UKF based on bearing-only observations is designed for the state estimation of multi-agents, rendering real-time and reliable movement control of the followers via the input-output feedback control law (Section 3) which gets rid of off-the-axis points. Properly designed simulations are given (Section 4) to demonstrate the validity of our approach. We offer our conclusions (Section 5). This section will formulate the \u201c1-N\u201d bearing-only observation model and give the notations used. As shown in Fig. 1, R1 represents the leader while R2 is a follower (for simplicity, we only show one here). The control inputs for agents are linear and angular velocities , , 1,2, \u2026., is the distance from the centroid of the leader to the centroid of the follower. is the view-angle from the y-axis of the follower to the centroid of the leader-robot. and are the orientations of the leader and the follower with respect to the world frame W , respectively, while is the relative orientation between the leader and the follower robot, i.e., . With reference to Fig. 1, the kinematic model of a 1-1(one leader and one follower) formation can be expressed as follows [ ]\u03a4 T\u03a4 1 1 1 ( , ) ( ) : ( ) ( ) = n f s u F s U S h s h s \u03d5 \u03b1 = = = = s y (1) where state vector T T , T , , input vector T, output vector T T, T and 1 1 1 1 1 1 cos 0 -cos 0 ( ) -sin / 0 sin / -1 0 1 0 -1 F s \u03b4 \u03d5 \u03b4 \u03c1 \u03d5 \u03c1 = . The kinematic model of the \u201c1-N\u201d (one leader and follower) formation can be readily retrieved as an extension of (1), ( ) ( ) TT T 1 1 1 +1 +1 T T 1 1 ( , ) , , ." ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001535_s1068366615020063-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001535_s1068366615020063-Figure4-1.png", "caption": "Fig. 4. Adopted pressure distribution in a single undulation sector of the end face gap.", "texts": [ " This result shows that, if cavitation is absent, the undula tion of the gap produces no hydrodynamic force. When a lubricant film breaks down in the cavita tion zone, the pressure p = ps.v.p. and the negative part of the hydrodynamic bearing capacity becomes lower in magnitude than its positive part. In this case, the seal has hydrodynamic bearing capacity. To make calculations more convenient, we will introduce a simplification by changing the shape of the cavitation zone (Fig. 2) to become rectangular (Fig. 4) with its radial boundaries rb1 and rb2, which coincide with the boundaries of the cavitation zone in its central part. Let us find the boundaries of the cavitation zone at \u03d5 = 1.5\u03c0/k. In this case, E = 6\u03bc\u03c9ek, h0 = hc. We will consider the expression for determining pressure (4) without allowing for inertia forces. At r = rb1 and r = rb2, the pressure p = ps.v.. ( ) ( ) ( ) ( ) ( ) ( )2 2 1 2 2 2 2 2 2 1 2 2 2 1 2 1 1 1 2 1 2 1 2 2 min min 2 2 2 12 22 2 1 21 1 ln 2 2 2 ln 3 1 1 16 2 ln . 2 ln r r r r p r r r W p p r r k h h e r rr r r r r rr r \u23a1 \u23a4\u2212\u2212\u23a2 \u23a5\u03c0 \u2212 \u23a3 \u23a6= + \u03c0 \u2212 \u23a7 \u23ab\u23aa \u23aa + \u03bc\u03c9 \u2212\u23a8 \u23ac +\u23aa \u23aa\u23a9 \u23ad \u2212\u23a1 \u23a4\u00d7 + \u2212 \u2212\u23a2 \u23a5\u23a3 \u23a6 stat dyn1 1 1 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000709_psc49016.2019.9081455-Figure19-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000709_psc49016.2019.9081455-Figure19-1.png", "caption": "Fig. 19. Flux leakages at the transverse and longitudinal edges [14]", "texts": [ " 18, LIMs are composed of on board armature windings and a conductive layer installed on the track and as the electrical energy needs to be transferred into the train, the electric connectors are necessary, limiting the speed of train. On the other hand as air gap of maglev trains, compared to rotational motors, is usually higher, the power factor and efficiency of these motors is normally very low [8]. Another difference between the radial and linear induction motors, is the presence of flux leakage at the edges of motor. These flux leakages occur at the transverse and longitudinal edges as shown in Fig. 19 [14]. It could be shown that in order to increase the powerfactor and efficiency by reducing the leakage flux of LIM, edges of the conductive track can be curved as illustrated in Fig. 20 [15]. Another way of reducing the end effects is by using the cylindrical design which is shown in Fig. 21 [16]. IV. SUSPENSION CONTROL IN MAGLEV TRAINS As electrodynamic suspension maintanes the air gap without needs for a control system, in this chapter, only the electromagnetic suspension is considered. The topology of the electromagnetic suspension is demonstrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000005_gtsd.2018.8595521-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000005_gtsd.2018.8595521-Figure5-1.png", "caption": "Figure 5. The yaw, pitch and roll angles.", "texts": [ " In which, r(t) is the desired process value or set-point (SP), y(t) is the measured process value (PV), e(t) is the error value which is the difference between the SP and the PV. The PID controller attempts to minimize the error over time by adjustment of a control variable u(t). The control function can be expressed mathematically as follows: u(t) = KP.e(t) + KI.0\u222bte(t\u2019)d(t\u2019) + KD.de(t)/dt (1) Next part presents PID controllers for the quadcopter balancing. The feedback are yaw, pitch and roll angle values calculated from the accelerometer sensor MPU 6050. The yaw, pitch and roll angles of the quadcopter are initialized as figure 5: Equation to calculate the angles from the gyroscope [2]: 0 sin cos 1 0 cos cos sin cos cos cos sin sin cos cos x y z t t t Where \u03a8, \u03b8, \u03a6 are the angles at the previous time (t-1). So, the real values of them at the current time are computed as below: ( ) ( 1) ( ) ( 1) ( ) ( 1) tt t t t t t t t t Based on x-axis, y-axis and z-axis accelerations measured from the accelerometer, the roll and pitch angles are computed [2]: cos cos cos cos sin 0 cos sin sin cos cos sin sin sin sin cos 0 cos sin cos sin cos sin cos sin cos cos 1 x y z A A A Thus, the actual values of roll and pitch angles measured by the accelerometer are calculated as below: arcsin cos yA Roll arcsin( )xPitch A In addition, in this platform, an ultrasonic sensor is used to maintain the altitude of the quadcopter in space" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure2-1.png", "caption": "Fig. 2. Blade coordinate system.", "texts": [], "surrounding_texts": [ "suffer from much more severe operational conditions compared to small bearings used for general industrial applications; in particular, they are subjected to greater external forces in each direction, very thin lubrication film thickness during operation, and so on. Furthermore, because they are installed on the wind turbine blade or a tower top of high altitude, it is costly and time consuming to repair or replace them in the field. Therefore, a laboratory-scale test is required for ensuring the performance of the pitch and the yaw bearings before application in the field. A test rig is developed to verify the performance of the pitch and the yaw bearings, including their fatigue life and static loading capacity. The test rig can reproduce actual operational conditions such as 6 degree of freedom (DOF) dynamic loadings and rotation of bearings for both directions. The mounting interfaces of the test rig are also the same as those used in the original environment, and various sizes of bearings can be applied by using a changeable adaptor. This high reproducibility of actual loading, driving, and mounting conditions simultaneously as well as applicability to wide size ranges are distinctively advantageous characteristics compared to previous test rigs. A structural analysis and preliminary friction torque test showed the suitability of the developed rig for use in pitch and yaw bearings of 2.0-3.0 MW class wind turbines.\nKeywords: Pitch bearing; Test rig; Wind turbine; Yaw bearing ----------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------\nThe pitch bearing and yaw bearing of a wind turbine are mechanical rotational elements that are essential to its safe and efficient operation [1]. The former enables relative rotational motion of the blade against the hub in a certain wind speed range, and the latter enables the tower to rotate against the nacelle to track the wind direction.\nBoth bearings should be capable of supporting a very heavy load of the order of several thousand kilonewtons and should have large inner and outer diameters of the order of several meters. Furthermore, they have rotational speeds of the order of a few RPM, which is quite low compared to industrial bearings having speeds of the order of several thousand RPM. The high load and low rotational speed result in a very severe lubrication environment at the rolling element contact surface, called as boundary lubrication [2]. These unique operational conditions should be accurately captured in the design stage, and the design should be verified through a laboratory-scale test. Robust design and verification tests are indispensable\nprocesses to ensure the reliability of pitch and yaw bearings. Several international standards [3, 4] and guidelines [5, 6] deal with the design of bearings, including methods for calculating the safety factor for static loadings and the endurance life for fatigue loadings. However, there exist no standards or guidelines regarding test items or procedures to ensure design robustness. Cumulative experience and know-how is required to establish reliable test procedures, for which purpose a test rig that can closely reproduce the actual operational environment of pitch and yaw bearings is first required.\nIn this study, we developed a new test rig for the pitch and yaw bearings of a wind turbine. This rig is unique compared to previous rigs in that it can apply all possible dynamic load components, including tilting moment, and faithfully reproduce actual operational environments including driving and mounting conditions. Furthermore, it can be applied to pitch and yaw bearings of various sizes and to the performance test of pitch and yaw drives. The applicability of the test rig to 2.0- 3.0 MW class wind turbines was shown through structural analysis and preliminary friction torque tests.\nOnly a few test rigs have been developed for testing large\n*Corresponding author. Tel.: +82 42 868 7994, Fax.: +82 42 868 7477 E-mail address: yjpark77@kimm.re.kr \u2020 Recommended by Editor Sung-Lim Ko \u00a9 KSME & Springer 2014", "wind turbine bearings. Among them, Astraios from Shaeffler Technologies is the best known one (Fig. 1) [7]. This rig comprises eight cylinders\u2014four radial and four axial\u2014that can be used to represent 6 degree of freedom (DOF) loads. Various types of bearings, including general slewing bearings and main bearings, can be tested at rotation speeds of 4-20 RPM through the drivetrain connection. The testable maximum outer diameter is 3.5 m. However, this rig is more suitable for testing main bearings than for testing pitch and yaw bearings; this is because the drive type used is a drivetrain that is suitable for main bearings. In contrast, pitch and yaw bearings respectively use a pitch and yaw drive comprising a motor, reduction gearbox, and pinion as a driving system. Consequently, this test rig cannot reflect the effect of the pitch and yaw drive on the bearing performance, and this may result in some discrepancies between the test results and the actual working performance.\nMTS\u2019s test rig can be used for all large wind turbine bearings including the main bearing and pitch and yaw bearings [8]. 6-DOF loadings and oscillatory motions can be simulated. Furthermore, bearings of various sizes can be applied by using changeable adaptors. However, the drive type of this rig, too, differs from that for pitch and yaw bearings, and full rotation motion is impossible if a cylinder is used as the driving unit.\nPSL\u2019s test rig for large wind turbine bearings was originally developed for testing only main bearings [9]. It can achieve a otational speed of 25-40 RPM and bear a large axial loading of up to 5,000 kN. However, it cannot be used to represent 6- DOF loadings because it does not have a radial direction loading system.\nFor wind turbine component design, the coordinate systems used include the blade coordinate system, chord coordinate system, hub coordinate system, rotor coordinate system, tower top coordinate system, and tower bottom coordinate system [5]. The blade coordinate system and tower top coordinate system are generally applied for pitch and yaw bearings, as shown in Figs. 2 and 3, respectively.\nThe loads experienced by pitch and yaw bearings under actual operational conditions can mainly be divided into extreme and fatigue loads [5]. Extreme loads are the most severe loads applied during the service life; these are static loads that are applied for a short time. Extreme loads are normally expressed as shown in Table 1, where \u03b3F denotes the safety factor for each load case, and the entire value was arbitrarily selected. Fatigue load is a repetitive load that is applied during the service life, and it is expressed as the load duration distribution (LDD) or damage equivalent load (DEL). LDD is a load spectrum that shows all load cases with their applied time. It is a static load with only a mean value, and in some cases, it may additionally also include oscillation amplitudes and oscillation speeds. DEL is an equivalent load that causes the same damage as that caused by all types of loads applied to the bearing during its service life. In general, the rain flow counting (RFC) method is used to obtain the DEL from unordered loading data, and DEL is a dynamic load that simultaneously has both an average and a range value. Furthermore, DEL includes the reference cycle, that is, the frequency of the loading as well as the S-N curve slope, to select the proper loading level according to the properties of the used material as shown in Table 2. The values in Table 2 were also arbitrarily selected.\nAs shown in Figs. 2 and 3, the loadings applied to the pitch and yaw bearings are expressed using six types of forces and moments: Fx, Fy, Fz, Mx, My, and Mz. Of these, Fz and Mz are 2-DOF axial loads, and Fx, Fy, Mx, and My are 4-DOF radial loads. Radial loads are often combined to obtain a single root mean square (RMS) value that is expressed in terms of the resulting transverse force, Fres, and resulting bending", "moment, Mres. General-use industrial bearings are designed by considering only force components because they are usually mounted on high-speed shafts that mainly perform power transmission, and they suffer from little shaft bending moment [3, 4]. However, pitch and yaw bearings in wind turbines support heavy interfacing structures that rotate at much lower speeds. Therefore, depending on the structural flexibility of the interfacing structures, these bearings may suffer large bending moments. In fact, the bending moment is known to have the greatest influence on the fatigue life of these bearings, and therefore, it should be carefully considered during their design stage. The various loading components applied to pitch and yaw bearings can be converted into a dynamic equivalent axial load, which is an axial stationary load that has the same fatigue effect as all loading components combined. The National Renewable Energy Laboratory (NREL), USA, devised an algebraic formula that could be used to calculate the dy-\nnamic equivalent axial load using bearing specs and loading information [6]. However, because bearing failure typically occurs at the interface of the maximum loaded ball and the raceway and it is impossible to decide the effect of each of many loading components on individual ball loading, it is necessary to test pitch and yaw bearings under actual load conditions to ensure their operational reliability.\nThe pitch and yaw bearings used in wind turbines have numerous bolt holes on their inner and outer ring, and they are respectively bolted to the hub and blade and the tower and nacelle during normal operation. To test these conditions, these bearings should be bolted to the corresponding frame.\nFurthermore, pitch and yaw bearings can be of internal or external teeth type, respectively shown in Figs. 4 and 5, depending on the location of the integrated gear teeth. Pitch bearings are typically of the internal teeth type, whereas yaw bearings can be of both types. Therefore, the test rig should be designed so as to be able to test both types of bearings. For internal teeth type pitch bearings, the outer ring is bolted to the hub and remains stationary, whereas the inner ring is bolted to the blade and rotates according to the wind speed. Therefore, the bearing\u2019s inner and outer rings should be divided into loaded and stationary parts to apply test loads to the loaded part while the stationary part is connected to the fixed supporting structure.\nThe test rig should be able to load and rotate the test bearing simultaneously. It is impossible to perform the two functions simultaneously using only a set of bearings because of the structural configuration of the bearings and the test rig. To solve this problem, by using two sets of the same bearings, two rings with gear teeth attached for both bearings are con-" ] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure15.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure15.5-1.png", "caption": "Figure 15.5 Strip heater", "texts": [ " Care must be taken not to overheat, as permanent damage to the material can result. Provided no permanent damage has been done, a shaped thermoplastic sheet will return to a flat sheet on the application of further heat. Simple bending is carried out by locally heating along the bend line, from both sides, until the material is pliable, using a strip heater. A strip heater can easily be constructed using a heating element inside a box structure, with the top made from a heat-resisting material. The top has a 5 mm wide slot along its centre, through which the heat passes, Fig.\u00a015.5. When the material is pliable, it can be located in a former and bent to the required angle, e.g. in making a splash guard for a lathe, Fig.\u00a015.6. The material can be removed from the former when the temperature drops to about 60 \u00b0C. Formers can be simply made from any convenient material such as wood. For volume production large-scale industrial machines are available which can have multiple bending areas and advanced heat control. Shapes other than simple bends can be carried out by heating the complete piece of material in an oven" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure1.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0002478_032026-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002478_032026-Figure1-1.png", "caption": "Figure. 1. Upgraded screw conveyor: A) a longitudinal section of the proposed screw conveyor; B) a section along A - A in figure 1 A.", "texts": [ " Based on the foregoing, we set a goal to find ways to improve the efficiency of cleaning cotton seeds before lintering and as the object of research we chose a screw conveyor, which is widely used in ginneries for transporting cotton seeds from saw gin to linter machines. Screw conveyors are simple in design, compact and reliable in operation [6,7]. To conduct experimental studies, we developed a pilot installation of a screw conveyor, where the lower half-cylindrical part of its trough was made in the form of a screening surface (mesh) for removing weed impurities emitted during transportation of cotton seeds (figure 1). The screw conveyor contains a groove 1, the lower part of which is in the form of a semi-cylinder, closed on top by a cover 2. The lower cylindrical part of the groove has openings 3, to remove debris emitted during transportation. The cover has an inlet 4 in the left part, and the groove 1 has an outlet 5 in the right part, a screw shaft 6 is installed inside the chute 1. To remove the removed weed impurities, an inclined tray 7 is installed at the bottom of the gutter. Bulk cargo-cotton seeds are fed into the groove 1 through the inlet 4 in the lid 2 and, when the screw 6 is rotated, is advanced by sliding along the groove 1, pushed by the working surface of the rotary screw 6 to the outlet 5" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003085_b978-0-08-099418-5.00009-3-Figure9.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003085_b978-0-08-099418-5.00009-3-Figure9.9-1.png", "caption": "Figure 9.9 The ring compression test.", "texts": [], "surrounding_texts": [ "Januszkiewicz and Sulek (1988) used the \u201crefusal technique\u201d to monitor the coefficient of friction necessary to initiate entry of the strip into the roll gap in a study of the effects of contaminants on the lubricating properties of lubricants. This approach makes use of the minimum coefficient of friction, required to initiate the rolling process. Recalling Eq. (4.1), the coefficient needed to allow entry is dependent only on the bite angle. At small reductions the bite angle is small and the required coefficient is also small. This fact is employed in the rolling process in which progressively smaller reductions are attempted in each pass. The bite angle at which entry is first successful is then reported as the coefficient of friction." ] }, { "image_filename": "designv11_22_0001286_s2095-7564(15)30296-8-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001286_s2095-7564(15)30296-8-Figure6-1.png", "caption": "Fig. 6 Time history of of acceleration and its time-frequency analysis ( V =50 km/h, L =10 mm)", "texts": [ " ( 16) gets the following result l /i K = - 2~ C = e-t (17) The relationship of coefficient u with time resolu tion ratio and frequency resolution ratio is constructed through Eq. ( 17) to determine the scale coefficient of FSWT. The FSWT is used to analyze the time-frequency characteristics of the axle box vertical acceleration sig nals under different conditions. When the running speed of vehicle is 50 km/h and the size of wheel flat is 10 mm, the time history of axle box vertical accel eration and its time-frequency analysis are shown in Fig. 6. When the running speed of vehicle is 110 km/h, and the size of wheel flat is 10 mm, the time history of axle box vertical acceleration and its time-frequency analysis are shown in Fig. 7. When the running speed of vehicle is 130 km/h, and the size of wheel flat is 40 mm, the time history of axle box vertical acceleration and its time-frequency analy sis are shown in Fig. 8. When the running speed of vehicle is 150 km/h and the size of wheel flat is 70 mm , the time history of axle box vertical acceleration and its time-frequency analysis are shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000549_0731684419888588-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000549_0731684419888588-Figure3-1.png", "caption": "Figure 3. The geometric parameters of the 2D cell and their effects on relative density. (a) Cell structure, (b) cell angle, (c) cell length, and (d) thickness ratio.", "texts": [ " The 3D cell is made of the 2D chiral cell, which looks like a cube. Every surface of the cube is a 2D cell, and thus the cell walls of the 2D cell are connected. Therefore, the mechanical properties are determined by the shape of the 2D cell. The mechanical properties of cellular structures are related to the relative density, so it is very important to work on the relative density before studying the axial compression performance of the CTCS structure. The structure of the 2D cell is shown in Figure 3(a), and t is the thickness of the cell wall, l is the length of the cell, and h is the angle between the cell wall and the cell diagonal. For cellular structure, therefore, the relative density is defined as follows qRD \u00bc qc qs (1) where qc is the effective density of the cellular structure and qs is the material density of the cell wall. Thus, the relative density can be written as follows qRD \u00bc Vwall Vcell (2) where Vwall is the volume of the cell wall and Vwell is the volume of the cell, which can be written as Vwall \u00bc 12 ffiffiffi 2 p lt2cosh\u00fe 6 ffiffiffi 2 p plt2sinh 12 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 p t5lsinh p , Vwell \u00bc l3. Therefore, the relative density is as follow qRD \u00bc 6 ffiffiffi 2 p a2 cosh\u00fe psinh\u00f0 \u00de 12 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 p a5sinh q (3) where a \u00bc t l is the thickness ratio. The effect of geometric parameters on the relative density is shown in Figure 3(b) to (d). On the one hand, with the increase of cell angle and thickness ratio, the relative density of cell increases, and the thickness ratio has a more obvious effect on relative density. On the other hand, with the increase of cell length, the relative density of cell decreases rapidly, but its range of change is small. Figure 4 is the deformation diagram of CTCS. When the compressive loading is applied in the vertical direction, the cellular structure has vertical compressive strain ey. At the same time, since the deformation of the 2D cell is not symmetric, the whole structure has a twist angle /, which presents the compression-twist effect" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002012_j.ifacol.2016.09.044-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002012_j.ifacol.2016.09.044-Figure1-1.png", "caption": "Fig. 1. The Qball-X4 quadrotor (Quanser (2013)).", "texts": [ " 20th IFAC Symposium on Automatic Control in Aerospace August 21-25, 2016. Sherbrooke, Quebec, Canada Copyright \u00a9 2016 IFAC 254 Simulation is implemented in Section V to illustrate the effectiveness of the designed controller. This study is finally concluded in Section VI. 2. DYNAMIC MODELLING AND CONTROL OBJECTIVES The coordinates of the quadrotor system\u2019s body frame {Ob, xb, yb, zb} centered at the center of gravity (CG) of the quadrotor, the global frame {Og, x, y, z}, thrusts, moments and gravity are represented in Fig. 1. Using Euler angles \u03d5 = [\u03c6, \u03b8, \u03c8]T and the rotational matrix R \u2208 SO(3) from the the body frame to the global frame, and following the Newton-Euler formalism, the dynamic model is derived based on applied forces F \u2208 R3 and moment M \u2208 R3 (Ko\u0308ksal et al (2015)), and is given by F = RFb = mp\u0308 and M = Jw\u0307 + w \u00d7 Jw (1) where R is the rotational matrix; Fb = [Fxb, Fyb, Fzb] T = [0, 0, \u22114 i=1 Ti] T is the applied force vector generated by actuators\u2019 thrust forces Ti, i = 1, 2, 3, 4, in the body frame; m is the total mass of the system; J\u03d5 = diag(J\u03c6, J\u03b8, J\u03c8) is the rotational inertia matrix in the body frame; w = [\u03c6\u0307, \u03b8\u0307, \u03c8\u0307]T is the angular velocity of Ob" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure11.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure11.4-1.png", "caption": "Figure 11.4 Top of column of vertical milling machine", "texts": [ " The capacity of these machines is identified by the size of the working surfaces of the table, the length of travel of the longitudinal, transverse and vertical movements, and the maximum distance from spindle to table surface on the horizontal model or from spindle to column on the vertical model. Figure 11.1 Horizontal milling machine CHAPTER Milling 11 D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 11 Milling 11 160 except that the spindle head is mounted at the top of the column, as shown in Fig.\u00a011.4. The column and base form the foundation of the complete machine. Both are made from cast iron, designed with thick sections to ensure complete rigidity and freedom from vibration. The base, upon which the column is mounted, is also the cutting-fluid reservoir and contains the pump to circulate the fluid to the cutting area. The column contains the spindle, accurately located in precision bearings. The spindle is driven through a gearbox from a vee-belt drive from the electric motor housed at the base of the column" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002060_j.wear.2016.11.002-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002060_j.wear.2016.11.002-Figure3-1.png", "caption": "Fig. 3. Schematic of the MRWAT apparatus.", "texts": [ " Modified abrasion test In order to measure specific energy, a modified rubber-wheel abrasion test (MRWAT) has been designed based on the standard dry sand/rubber wheel abrasion test (ASTM-G65) [9]. The main advantage of the modified test is its ability to better replicate actual abrasive field conditions experienced in an oil sands mining environment. A total of 85 test runs were performed; 37 samples for Al 61 and 16 each for Al 63, mild steel A36, and stainless steel 17-4SS. A schematic diagram of the experimental setup of the MRWAT test is shown in Fig. 3. This apparatus is similar to the standard ASTM G65 setup but with two major differences: the abrasive medium used and the wheel. Oil sand has zero cohesion but high friction angle [16]. The quartz particles in oil sand vary in shape from angular to rounded, generally described as being sub-angular. The sand used in the MRWAT is in-situ oil sand (bitumen stripped) rather than the standard AFS 50/70 test sand which consists of rounded quartz grains and is generally used in ASTM G65. Secondly; the ASTM G65 procedure specifies a rubber wheel of with a 22" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure4.15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure4.15-1.png", "caption": "Figure 4.15 Right-angled bracket", "texts": [ " O b1 1 2 3 4 5 6 7 8 9 10 11 12 2 3 4 5 6 789 10 11 12a Circumference \u03c0d Figure 4.11 Development of cone O c ba Circumference \u03c0d Figure 4.12 Development of part cone D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 4 Sheet-metal operations 59 To find the development length on the flat sheet, it is necessary to find the length of the mean line by calculating the lengths of the flat portions and the bends separately. The stretched-out length of the bend is called the bend allowance and for a by 1.57 (i.e. \u03c0/2). Figure 4.15 shows a right-angled bracket made from 1 mm thick material. To obtain the development, first find length ab = 60 \u2212 inside radius \u2212 metal thickness = 60 \u2212 2 \u2212 1 = 57 mm next length cd = 80 \u2212 2 \u2212 1 = 77 mm finally length bc = mean radius \u00d7 1.57 = (inside radius + 1 2 metal thickness) \u00d7 1.57 = (2 + 0.5) \u00d7 1.57 = 2.5 \u00d7 1.57 = 3.9, say 4 mm \u2234 total length of development = 57 + 77 + 4 = 138 mm 59 57 77 Bend line Bend allowance 4 da b c Figure 4.16 Development of right-angled bracket It can be seen that the development of the bracket shown in Fig. 4.15 is made up of a 57 mm straight length plus a bend allowance of 4 mm plus a further straight length of 77 mm, as shown in Fig. 4.16. The bend is half way across the bend allowance, and therefore the bend line must be 57 + 2 = 59 mm from one edge. Review questions 1. Why is it necessary to overbend material during a bending operation? 2. Under what circumstances would it be more appropriate to use a folding machine rather than a vice when bending sheet metal? 3. Why is it necessary to calculate the developed length of sheet metal components using the mean line" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000267_1.g003407-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000267_1.g003407-Figure1-1.png", "caption": "Fig. 1 Dynamic soaring cycle of bend type.", "texts": [ " Emphasis is placed on the manner in which the energy gain from thewind is achieved so that flight without an energy effort by a flying object is possible. Thus, it will be shown that the kinetic energy concepts under consideration are of equal significance for describing energy harvesting. As a consequence, the concepts of kinetic energy related to airspeed and to inertial speed are equivalent. Dynamic soaring is a non-powered flight mode that consists of periodically repeated cycles one of which is schematically presented in Fig. 1 and 2 respectively. There are two cycle forms that may be regarded as basic types of dynamic soaring. One is S shaped andmay be termed as bend type (Fig. 1), characterized by the fact that no Received 25 October 2017; revision received 31 July 2018; accepted for publication 10 January 2019; published online 28 June 2019. Copyright \u00a9 2019 by Gottfried Sachs. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission. All requests for copying and permission to reprint should be submitted to CCC at www.copyright.com; employ the eISSN 1533-3884 to initiate your request. See also AIAA Rights and Permissions www.aiaa.org/randp", " That work scenario holds in the lower curve of a dynamic soaring cycle (flight phase 4). Because the wind speed is small here (as supposed for Fig. 5), the resulting negative work is also small. The described energy characteristics have been determined for an exemplary case in a computational treatment of dynamic soaring. The performance goal was to minimize the required shear wind strength necessary for energy-neutral dynamic soaring (details are given in Appendix B). The addressed kind of dynamic soaring cycle is of type 1 (according to Fig. 1), the flight path of which is presented in Fig. C1 in Appendix C. The case dealt with is regarded to be representative for describing the energy characteristics under consideration. Results are presented in Fig. 6, which shows the time histories of the kinetic, potential and total energy as well as the inertial speed and the altitude. In the upper part of Fig. 6, the time history of the total energy Etot;in is plotted. The flight phase in which the energy gain from the wind is accomplished is highlighted by gray shading" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000338_rpj-11-2018-0291-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000338_rpj-11-2018-0291-Figure8-1.png", "caption": "Figure 8 Error caused by circular arc transition method", "texts": [ " By limiting the maximum normal acceleration of small arcs, the speed of the corners can be restricted and the accuracy effect is considered. According to the formula of kinematics, the normal acceleration of circularmotion is: aset \u00bc Vlimit2 2 r (19) Here aset is the maximum acceleration specified by the user, and Vlimit2 is the maximum allowable velocity at the inflection point calculated from the maximum acceleration. r is the radius of curvature of the arc transitionmethod for two line segments. As shown in Figure 8, Ln-1 and Ln are adjacent lines in the trajectory, and d is the maximum error produced by the arc transition method. u is the angle between Ln-1 and Ln, and R is the radius of the arc. Cn is the intersection point between Ln\u20131 and Ln. The function of these parameters can be written as the following formula: sin p u 2 \u00bc r d 1 r (20) Comprehensive formula 19, 20 can be obtained: Vlimit2 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi aset d cos u 2 1 cos u 2 s (21) Integrating Equation (9), Equation (11), Equation (12) and 16, Equation (21), the optimal linked feedrate is shown as Equation (22)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001279_s0263574714000149-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001279_s0263574714000149-Figure5-1.png", "caption": "Fig. 5. A different KCB discretization of the mechanism.", "texts": [ " For the sample mechanism the mobility is therefore M = \u2211 fi \u2212 \u2211 rk = 5 \u2212 (2 + 2) = 1, (7) which means the kinematic formulation includes 5 equations (4 constraints plus one driving motion) while the dynamic formulation includes a DAE system with 9 equations (5 differential and 4 algebraic equations). The discretization of the mechanism in KCBs is not unique but can be performed in various ways; this now influences the analytical models representing the kinematics and dynamics. For example, the open loops could be created not by opening joints but by dividing bodies as shown in Fig. 5 for the same planar mechanism. In this case, for each loop closure 3 constraints equations are needed to express the assembly conditions for an individually discretized body. The mobility formula is the same as previously, the result identical (M = 1), but the terms of the formula have different values: M = \u2211 fi \u2212 \u2211 rk = 7 \u2212 (3 + 3) = 1. Despite the identical result of the mobility calculation, the terms of the relation suggest different numbers of equations for the kinematic and dynamic formulation, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000011_icarcv.2018.8581093-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000011_icarcv.2018.8581093-Figure4-1.png", "caption": "Fig. 4 The four-UAV motion trajectories around the stationary target", "texts": [ "1 The relative distance of between UAV and stationary target t(x10s) R e la ti v e S p a c in g E rr o r( x 1 0 m ) Fig.2 The relative distance error between two UAV around the stationary t(x10s) V e h ic le A ir s p e e d ( x 1 0 0 m /s ) Fig.3 The airspeed command of between the spacing and maintaining circular orbit around The Fig. 1 adopts the three UAVs formation as a control object and presents the relative distance error between the UAV and the target in the process of tracking motion target, which shows a trend of tending to level state after rapid decline. As can be seen from the figure 4, the UAV formation consists of one leader and two followers and is rapidly close to the target. From 1s to 15s, the leader of formation is in front of the follower, and the forward velocity is larger than two followers. In addition, the UAV tracks target with the fastest velocity and most optimized path before the UAV is close to the target; when the UAV is very close to the target, the UAV formation can achieve the purpose of cooperative tracking with a fixed geometric shape. The Figs. 2 and 3 adopt the two UAVs as a control object. As can be seen from the figure 3, it shows that the relative distance error between two UAVs around the stationary target, and presents a tendency to decrease and then increase. In the process of the tracking the target with the shortest time, the multi-UAV will be affected by air resistance because the wingtip will generate the vortex effect. From figure 4, it presents that the airspeed command of between the spacing and maintaining circular orbit around the motion target, and shows a trend of the cooperative formation flight, but there is a certain delaying phenomenon. In the process of the cooperative tracking controller, each UAV is equipped with a control command sensor. Especially, there is a phenomenon that the received information is deviated in the process of high-speed flight. Fig. 4 shows the four UAVs\u2019 trajectories around the stationary target. Initially, the UAV has multiple headings and loiter circles, all of which converge to the desired heading along the Lyapunov guidance vector field. When all UAVs reach a loiter circular orbit with the prescribed distance, the variable airspeed controllers can achieve the desired angular spacing. In this paper, we design the control laws and prove the stability of tracking motion target based on the Lyapunov guidance field, which is only theoretical research" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003001_j.engfailanal.2020.105195-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003001_j.engfailanal.2020.105195-Figure6-1.png", "caption": "Fig. 6. Geometry of the unnotched specimen for a monotonic tensile test made of 25CrMo4 steel, units in mm.", "texts": [ " Table 1 shows the percentage composition of elements compared to the standard requirements [26]. Values correspond to the standard. The mechanical properties were determined using Instron ElectroPuls E3000 - a dynamic and fatigue testing machine with electrodynamic drive in accordance with ISO 6892 [27]. Static tensile tests were carried out using a force gauge with \u00b1 5 kN measuring range and 2620 Instron axial extensometer with a base unit of 12.5 mm and the elongation range of \u00b1 5 mm. The tests were carried out on unnotched specimens (Fig. 6) taken from the top tube. The value of the initial elastic modulus E is obtained by a linear regression least-squares estimate of the slope in the proportional elastic region. Table 2 shows the average values (three specimens) for the mechanical properties (modulus of elasticity E, tensile strength Rm, yield strength Re, reduction of area Z, longitudinal elongation A). The test program involved the determination of fatigue properties of top tube specimens. High-cycle fatigue tests under axial loading were performed in accordance with standards [28,29]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003289_j.chroma.2021.461925-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003289_j.chroma.2021.461925-Figure2-1.png", "caption": "Fig. 2. Construction of the pump head module. The left-hand part (outer diameter 38 mm) contained the seal with an internal back-up ring and was connected through a female VICI Valco fitting to 1/16 inch PEEK capillary tubing. Left and right PEEK parts were assembled as indicated by the arrows using two bolts and nuts.", "texts": [ " Pump head We developed a PEEK pump head module ( Figs. 1 , 2 ) that as machined by Microtherm B.V. (Oudkarspel, The Netherlands). he outer diameter of the left-hand part (38 mm) as shown in ig. 2 was such that the module fitted into a standard caulk gun rom a local hardware store ( Fig. 1 ). A DS 119 PTFE seal (6 mm x 2 mm, 6 mm long) was obtained from Eriks (Alkmaar, The Netherands). A stainless steel (SS 316) piston with a diameter of 6 mm as generating the high pressure when pushing the piston to the eft ( Fig. 2 ). The piston was guided by a Teflon slider within the haft for positioning. A side branch was welded to the piston for etter grasp when pulling it for filling with new eluent. .2.3. Pulse dampener In order to reduce the fall in pressure during the running of chromatogram, a pulse-dampener with a PEEK body was deigned by us and machined at the Technics Campus Den Helder, he Netherlands. It consists of two PEEK disks of 24 mm (upper isk) and 10 mm thick (lower disk) that are pressed together by welve M5 bolts (see Figs", " A VICI Valco micro-injection valve model Cheminert C4-1004-.5 with stator bore of 0.25 mm and an nternal loop of 0.5 \u03bcL) was used in cases where an exact and fixed njection volume was needed, e.g. when the number electrons per xidized molecule was calculated during oxidation at different flow ates. . Results and discussion .1. Functioning of the pump module and the pulse dampener A working pressure of 100 bar was used most of the time. or building up this pressure, manual compression of one cyliner volume of the pump module (see Fig. 2 ) was sufficient. Howver, by strong squeezing of the caulk gun a pressure of 170 bar ould also be reached. At the start we injected approximately 20 L of the propanol/nitric acid mixture (for details, see above) in rder to abolish phase collapse at zero flow rate and for removing ighly retained compounds of previous injections. To avoid polluion of the electrode the column was uncoupled from the detecor during this procedure. The best time for this was late in the vening, since the following morning a series of chromatograms ould be started without losing too much time for stabilization" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003114_tmag.2015.2438872-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003114_tmag.2015.2438872-Figure1-1.png", "caption": "Fig. 1. Slinky-laminated stator core.", "texts": [ " Index Terms\u2014 Compressive stress, iron loss, magnetic property, slinky lamination, strain, tensile stress, washing machine. I. INTRODUCTION THE manufacturing of the magnetic parts of the electrical machines requires several industrial processes. These may have effects on the magnetic properties of the material, especially in terms of iron loss [1]\u2013[3]. With the development of manufacturing techniques and production machines, the slinky-lamination method is now being used to significantly reduce the material waste [4]. For this process, a long iron band is rolled up spirally and bonded, as shown in Fig. 1 [1]. Thus, not only compressive stress but tensile stress and strain occur in the core. As a result, the magnetic properties deteriorate, and it becomes hard to estimate the deteriorated properties of the slinky-laminated material. There have been several studies on the various manufacturing effects on the magnetic properties [1]\u2013[3]. In these studies, the effects of mechanical stress caused by wire cutting, shrink fitting, pressing, and the punching process were examined through the experimental methods" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001332_01431161.2014.967043-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001332_01431161.2014.967043-Figure1-1.png", "caption": "Figure 1. Forms of the investigated nozzles: (a) long nozzle and (b) short nozzle.", "texts": [ " The nozzles for the coaxial feed of powder materials are attachments for producing a stable gas-powder jet with the required characteristics. Different operations use nozzles of different shape: with an acute angle at the tip, with a large angle at the tip, with a small and large beam diameter in the treatment plane.1 Therefore, to optimize the shape of the adjustable nozzle for the coaxial feed of powder material, mathematical modelling was used for designing two variants of the nozzles in the form of a pair of attachments (external/internal), in order to determine the optimum combination of the attachments (Figure 1). The first variant has the cone angle at the tip of 508, the second 658. Using nozzles with these dimensions, the cladding zone can be displaced from the outlet of the tool by at least 10mm, and the diameter of the spot in the zone can be increased. To determine the optimum dimensions of the nozzle cone, the movement of the powder particles in the channel to the cladding surface was investigated by mathematical modelling. The aim of mathematical modelling was to verify the efficiency of both designs of the nozzle in cladding in the conditions of actual acting factors of the process and study the effect of the parameters on the gaspowder flow" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003350_s00202-020-01174-5-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003350_s00202-020-01174-5-Figure2-1.png", "caption": "Fig. 2 Machine structure", "texts": [ " The looping of the stator winding star point at the middle of the fourth switch leg of the inverter avoids the suppression of the generated third-harmonic currents in the stator winding star point. On the rotor side, an additional harmonic winding is wound alongside the conventional field winding. The harmonic winding is connected at the field winding through a rotary bridge rectifier mounted on the rotor\u2019s periphery. The machine system is based on a semi-open winding. The machine structure consists of 42 slots and 1 3 four poles, as shown in Fig.\u00a02. The relevant specifications are listed in Table\u00a01. The machine\u2019s rotor field winding is excited through a spatial third-harmonic magneto-motive force (MMF) component. The four-leg inverter and the semi-closed winding configuration create a path that allows the flow of the current of the third-harmonic component in the stator winding. The fundamental and third-harmonic currents in the stator winding produce a composite MMF in the machine air gap. The composite MMF contains an additional zero-sequence third-harmonic MMF alongside the positive-sequence fundamental MMF" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000853_978-3-642-40066-7_10-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000853_978-3-642-40066-7_10-Figure1-1.png", "caption": "Fig. 1 Examples of three-dimensional wireless sensor networks: a underwater sensor network, b underground sensor network, and c airborne sensor network", "texts": [ " In facing up to these challenges, there have been new network protocols and algorithms specifically designed for 3D wireless sensor networks using geometric approaches by exploring rich geometric properties of sensor networks. In this chapter, we will review the most recent advances in 3D wireless sensor networks with a focus on geometric approaches for two particular sensor network problems: 3D topology control and 3D geographic routing. Three-dimensional wireless sensor networks have a variety of applications. Figure 1 shows three possible applications of 3D wireless sensor networks. Underwater sensor network: Underwater acoustic sensor network (UWSN) [3] or ocean sensor network [98], shown in Fig. 1a, can find applications in oceanographic data collection, pollution monitoring, offshore exploration, disaster prevention, assisted navigation, and tactical surveillance applications. Three-dimensional UWSN is used to detect and observe phenomena that cannot be adequately observed by means of ocean bottom sensor nodes (in a 2D sensor network), i.e., to perform cooperative sampling of the 3D ocean environment. In 3D UWSN, sensor nodes float at different depths to enable the exploration of natural undersea resources and gathering of scientific data in collaborative monitoring missions", " Besides the 3D deployment, UWSN is also significantly different from terrestrial sensor networks: (1) acoustic channels used under water feature long propagation delays and low bandwidth; (2) underwater sensor nodes may move with water, thus introducing passive mobility. Underground sensor network: Underground sensor network (UGSN) [4] can be used to monitor a variety of conditions, such as soil properties for agricultural applications, integrity of belowground infrastructures for plumbing, or toxic substances for environmental monitoring. For example, agriculture can use underground sensors to monitor soil conditions such as water and mineral content [21]. Wireless sensors can also be used to monitor the underground tunnels in coal mines [52] as shown in Fig. 1b. These tunnels are usually long and narrow and distributed in 3D, with lengths of tens of kilometers and widths of several meters. A full-scale monitoring of the tunnel environment (including the amount of gas, water, and dust) has been a crucial task to ensure safe working conditions in coal mines. Airborne sensor network: Unmanned air vehicles (UAVs) has been proposed to be used as mobile, adaptive communication backbones for ground-based sensor networks [75, 86]. Figure 1c illustrates that multiple UAVs can serve as an airborne relay for a ground-based sensor networks. The UAVs and the sensor nodes together form a 3D airborne sensor network to support the military applications. The UAVs can also provide communication connectivity to sensors that cannot communicate with each other because of terrain, distance, or other geographical constraints. Moreover, the UAVs themselves can have sensing capacity and form a pure airborne sensor network [6]. Besides examples above, 3D wireless sensor networks can also be found useful in many other applications, such as a large 3D space network for space explorations [33] or a small 3D sensor network in a multi-floor building for structure monitoring [106]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003684_j.matpr.2021.05.466-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003684_j.matpr.2021.05.466-Figure17-1.png", "caption": "Fig. 17. Longitudinal Stress in Elliptical head cylinder.", "texts": [], "surrounding_texts": [ "The dimension was taken according to ASME standard section VIII, Division I, ASME Section VIII, Division I, UG-27 and calculations (equation (1) to (5)). Diameter \u2013 250 mm Length \u2013 420 mm Thickness \u2013 1.5 mm Figs. 8, 9, 10, and 11 demonstrate the 3D model of various heads cylinder. After Modeling, FEA was done which is explained below section." ] }, { "image_filename": "designv11_22_0002854_j.engfailanal.2020.105028-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002854_j.engfailanal.2020.105028-Figure8-1.png", "caption": "Fig. 8. Stress and deformation of connecting rod under the condition of thermo mechanical coupling.", "texts": [ " Zhang Engineering Failure Analysis xxx (xxxx) xxx According to the maximum data obtained from the simulation, the deformation and stress of the connecting rod under the maximum stress is analyzed. The maximum peak pressure is put to the surface at the small end of the rod and the various constraints is set as discussed earlier. After putting temperature field and stress load, the maximum pressure stress is 235 MPa, the yield strength of 7075-T6 aluminum alloy is 417 MPa under 125 \u2103, and safety factor is 1.77. The maximum deformation occurs at the top of the connecting rod, and the maximum deformation value is 0.13 mm, as shown in Fig. 8. Z. Pan and Y. Zhang Engineering Failure Analysis xxx (xxxx) xxx Under the action of thermo mechanical coupling, the safety factor of the weak area of the connecting rod is 1.77 which is more than 1.3, which meets the design requirement. But the piston of the engine does reciprocating movement for a long time, and the connecting rod is affected by alternating stress, resulting in high cycle fatigue. According to the maximum running speed of the engine 2700 rpm, the certification test time is 150 h" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002131_ecc.2016.7810331-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002131_ecc.2016.7810331-Figure1-1.png", "caption": "Fig. 1: 3D CAD model of the New Quadrotor Manipulation System", "texts": [ " However, the nonlinear control methods make the MPC to be more complex and has high computational cost which is not suitable to our robotic system. This paper is organized as follows: In section II, the considered robotic system is described, and kinematic and dynamic analysis are reviewed. The control problem to solve is formulated and the DOb and MPC approaches are described in section III. In section IV, simulation results using MATLAB/SIMULINK are presented. Finally, the main contributions are concluded in section V. 3D CAD model of the proposed system is shown in Fig. 1. The system consists mainly of two parts; the quadrotor and the manipulator. Fig. 2 presents a sketch of the proposed system with the relevant frames which indicates the unique topology that permits the end-effector to achieve arbitrary pose. The frames satisfy the Denavit-Hartenberg (DH) convention. The manipulator has two revolute joints. The axis of the first revolute joint (z0), that is fixed to the quadrotor, is parallel to the body x-axis of the quadrotor (see Fig. 2). The axis of the second joint (z1) is perpendicular to the axis of the first joint and will be parallel to the body y-axis of quadrotor at home (extended) configuration" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000514_j.mechmachtheory.2019.103667-Figure21-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000514_j.mechmachtheory.2019.103667-Figure21-1.png", "caption": "Fig. 21. Schematic representation of the deviation of the shaper axis during generation of the face gear to achieve the desired deviation \u03b4max on the face-gear tooth surface.", "texts": [ " 20 shows the results of comparison of the compensated geometry of the face-gear tooth surface after the third iteration of compensation (see Table 3) with respect to the standard geometry achieved with the generation of the face gear by a shaper with 26 teeth and no errors of alignment compensated. Notice that the deviations obtained after compensation of errors of alignment caused by shaft deflections show a longitudinal deviation pattern with a maximum of 140 \u03bcm at the inner radius of the face-gear tooth surface. It is equivalent to the application of a small left hand helix angle to the face-gear tooth surfaces. Fig. 21 shows the schematic representation of a deviated shaper for generation of the face gear in order to achieve the desired deviation \u03b4max on the face-gear tooth surfaces. This deviation is equivalent to the compensation of only an error E of minimum distance between the shafts of the shaper and the face gear. If the face width of the face gear is denoted by Fw2 and the required deviation of the face-gear tooth surface is \u03b4max at the inner radius (point A2 in Fig. 21), the shaft of the shaper has to be rotated around point A1 of an angle \u03bb given by \u03bb = arctan \u03b4max F (29) w2 Point A1 is located at the outer radius of the face gear denoted by L2. Therefore, the applied error E, as shown in Fig. 21, is easily determined as E = L2 tan \u03bb = L2 \u03b4max Fw2 (30) Considering the data shown in Table 1 for the face gear and a desired deviation at the inner radius of the face gear of 140 \u03bcm as shown in Fig. 20, the error E to be compensated is of 0.89 mm. For the convenience of identifying different designs, the reference design with all errors compensated, that is to say the face gear with compensated geometry for E = 1.2923 mm, A1 = \u22120.0188 mm, A2 = 0.4917 mm, and \u03b3 = \u22120.0304\u25e6, will be referred to as Case 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002953_0954406220974052-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002953_0954406220974052-Figure1-1.png", "caption": "Figure 1. The virtual prototype of HFFD-6.", "texts": [ " \u2018The gravity compensation approach\u2019 section presents the active and passive combined gravity compensation approach, and both the passive and active compensation approaches are evaluated by simulation. And then the experiments are presented in the penultimate section with the discussion on results. The final section concludes the works. In this work, the force compensation is conducted for a hybrid 6-DOF force feedback device\uff08HFFD-6), which composes of a parallel mechanism and a serial one. As shown in Figure 1, the parallel mechanism is a modified Delta parallel mechanism with 3- DOF in translational directions. While the serial mechanism has 3 rotational DOF. The parallel and serial structures can be decoupled; thus, the kinematic analysis can be obtained respectively. From Figures 1 and, 2 the parallel mechanism is a horizontally-mounted modified Delta mechanism. It composes a fixed base, a moving platform, three active linkages, and three passive parallelogram linkages. Three identical motors are mounted at the back of the fixed base" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003349_s00170-021-06813-0-Figure19-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003349_s00170-021-06813-0-Figure19-1.png", "caption": "Fig. 19 Thermal contact areas of 3-piece conventionally packaged cold plate highlighted in dark yellow. a Contact area between gauge 20 plate and block 1. b Contact area between gauge 20 plate and ceramic tube. c Contact area between interface of blocks 1 and 2. d Contact area between gauge 20 plate and block 3. e Contact area between gauge 20 plate and ceramic tube. f Contact area between interface of blocks 2 and 3", "texts": [ " Rc \u00bc LWhc\u00bd \u22121 \u00bc 1:25KsLW m sig P Hc :95 \u22121 \u00f07\u00de where, Rc is contact resistance hc is joint contact conductance Ks is harmonic mean thermal conductivity L is the length that the two surfaces have in common W is the width that the two surfaces have in common m is effective mean absolute asperity slope sig is effective RMS surface roughness P is contact pressure Hc is the surface microhardness of the softer of the two contacting solids And temperature difference due to contact resistance can be calculated as \u0394T \u00bc q LWhc\u00f0 \u00de \u00f08\u00de where, \u0394T is temperature difference q is applied heat load The contact surfaces of 3-piece cold plate packaged by conventional techniques are shown in Fig. 19. The joint contact conductance, contact resistance, and temperature difference for each interface are calculated per above formulas, and the results are tabulated in Tables 3 and 5. To validate the experimental results, temperature differences from above tables are compared to corresponding Table 5 Comparison of experimental and theoretical temperature differences at interfaces due to contact resistance of 3-piece conventionally packaged cold plate Experimental vs theoretical \u0394T due to contact resistances of blocks 1, 2, embedded ceramic tube, and gauge 20 plate \u0394T due to contact resistances of blocks 2, 3, embedded ceramic tube, and gauge 20 plate Total \u0394T due to all contact resistances for 3-piece cold plate 3-piece cold plate and (margin of error) \u0394T1=(TC1\u2212 TC3)conv\u2212 (TC1\u2212TC3)in situ from experiment (\u00b0C) \u0394T2 Cumulative \u0394T shown in Fig. 19 a, b, and c from theoretical formula (\u00b0C) \u0394T3= (TC4conv\u2212 TCpc)\u2212 (TC4insitu\u2212 TCpc) =(TC4conv\u2212 TC4insitu) from experiment (\u00b0C) \u0394T4 Cumulative \u0394T shown in Fig. 19 d, e, and f from theoretical formula (\u00b0C) \u0394T= \u0394T1+ \u0394T3 from experiment (\u00b0C) \u0394T= \u0394T2+\u0394T4 from theoretical formula (\u00b0C) BN-G20C-CSS 6.64 5.91 4.37 4.16 11.01 10.07 (Margin of error for BN-G20C-CSS) 10.9% 4.8% 8.5% BN-G20S-CSS 6.26 5.74 4.32 3.97 10.58 9.71 (Margin of error for BN-G20S-CSS) 8.3% 8.1% 8.2% temperature differences between conventional and AM cold plate from experiment (difference of thermocouple readings placed at both sides of contact interfaces). Therefore, this comparison would reveal temperature difference just due to contact resistance between pairs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000861_j.phpro.2015.06.044-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000861_j.phpro.2015.06.044-Figure1-1.png", "caption": "Fig. 1. Linear motor and flexure bearing assembly Fig. 2. Flexure bearings (a) Spiral; (b) Linear", "texts": [ " Keywords:cryocooler; linear compressor; flexure bearing; FE analysis; strain and stiffness measurement Lubrication is a common problem in miniature cryocoolers which contaminating the regenerators. Small capacity cryocoolers for space borne application generally use linear electromagnetic drives for long life, maintenance free operation and high reliability. Flexure bearings are used for dynamic support in linear motor compressors. A typical flexure supported electromagnetic drive system is shown in Fig. 1. For eliminating rubbing contact between the piston and the cylinder wall, flexure bearing is used to support the piston inside the cylinder. A * Corresponding author: E-mail address: marutikhot@gmail.com \u00a9 2015 The Authors. Published by Elsevier B.V. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/). Peer-review under responsibility of the organizing committee of ICEC 25-ICMC 2014 radial clearance gap of 10 to 20 \u03bcm between piston and cylinder provides the necessary flow impedance to serve as dynamic seal" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003519_s10846-021-01359-5-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003519_s10846-021-01359-5-Figure10-1.png", "caption": "Fig. 10 Various environmental cases under which the collaborative manipulation is formulated upon. Case I, is an obstacle-free environment, whereas cases II, III and IV are obstacle containing environment. Case IV is a combination of Case II and III", "texts": [ " Assumption 4 Obstacles Oi are placed in orientation such that they are detected by the RF , if there is a need to perform obstacle avoidance. Assumption 5 Wheel slip, either in RL or RF , has no effect on the overall maneuverability of the system. Assumption 6 Payload length is assumed to be > 4Lx and has width in the range of the clamping width. In order to ensure proper functioning of the collaborative strategy, each working case must be analyzed beforehand. With two robots RL and RF , the following scenarios are encountered while navigating an environment. CASE (I), as shown in Fig. 10a, presents two robots, RL and RF , in an obstacle-free environment. The (Vxf , Vyf ) and (Vxl , Vyl) represent the follower and leader robot\u2019s velocity along the robot\u2019s XR and YR coordinate axis, respectively. During the RL\u2019s motion, a force \u2212\u2192 F is produced on RF , whose magnitude directly relates to displacement\u201d and the direction is given by \u201c\u03b8\u201d. (\u03c9f 1, \u03c9f 2, \u03c9f 3, \u03c9f 4) represents the wheel velocity in rad/s for respective wheels, as shown. CASE (II) is a simple obstacle containing environment. It contains an object on the other side of the RL. For instance, during initialization of F from \u03b8 direction, an obstacle is present in the -\u03b8 side of RL as shown in Fig. 10b. CASE (III) is also an obstacle containing environment. It contains an object on the same side of the RL, contrary to CASE II. For instance, during initialization of F from \u03b8 direction, an obstacle is present in the \u03b8 side of the RL, as shown in Fig. 10c. CASE (IV) is a full-edged obstacle-clustered environment with a combination of CASE II and CASE III, as shown in Fig. 10d. In this case, at a particular interval of time, CASE IV is categorized into either CASE II or III and handled as such consequently. Figure 11 describes the coordination architecture between the leader and the follower. A user controls the leader robot where a payload is clamped on it as well as on the follower robot in such a way that there is no relative motion between any of the clamps and the payload. Whenever the leader robot moves, the payload moves along with it, producing forces on the follower robot", " Finally, ([Vxf , Vyf , \u03c9z]) are converted to (\u03c9f 1, \u03c9f 2, \u03c9f 3, \u03c9f 4) through an Inverse kinematic Equation 2, this contains the individual rpm required by each wheel to achieve the desired outcome and is hence uploaded to the respective motors. Algorithm 1 Updating (RF )Follower Robot\u2019s position. function UpdateManipulation(d, \u03b8, UL, UR) IF \u03b8 > 0 AND UL==TRUE OR \u03b8 < 0 AND UR==TRUE THEN Figure 12 and Algorithm 1 describes the flow of command taking place in the follower robot with the attributes driven from leader robot and its local environment. The follower robot\u2019s (RF ) algo.1. deployed in the micro-controller as shown in Fig. 12 is designed based on the cases mentioned in Fig. 10. The values received by the left ultrasonic sensor \u201cUL\u201d, right ultrasonic sensor \u201cUR\u201d, linear potentiometer \u201cd\u201d and rotary potentiometer \u201c\u03b8\u201d are input parameters for this algorithm. During cases CASES I and II, the robot moves along \u2212\u2192 F using Eq. 2. This is so because \u201cUL\u201d and \u201cUR\u201d have no impact on the overall trajectory. Taking case II for instance in Fig. 10b, an obstacle present in another direction where RF is moving along \u2212\u2192 F does not hamper its motion. On the other hand, CASE case III, where an obstacle is present in the vicinity of \u2212\u2192 F (either left or right), affects the motion of the RF . In this case, the algorithm sets the motion at \u201c-\u03b8\u201d direction to perform obstacle avoidance till the instance when there is no obstacle according to CASE case III (either left or right). CASE Case IV being a combination of CASE Case II and CASE Case III scenarios, it uses each of the conditions and their corresponding output to tackle such an obstacle-clustered environment" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001352_s0025654414010026-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001352_s0025654414010026-Figure3-1.png", "caption": "Fig. 3.", "texts": [ " Only the counterweight mass of 3000 kg located at a distance of 1.4 m from the crank axis, the oil column mass of 3000 kg lifted per one cycle, and the rod mass of 3000 kg are taken into account. The crank rotates counterclockwise at an angular velocity 0.534 rad/s. The energy losses in the actuating motor are taken into account according to (4.3), where all coefficients are zero except for \u03b12. The graph of the torque M (N\u00b7m) against the driving crank \u03d5 (rad) for the case in which the friction forces on the driving crank are neglected is given by curve 1 in Fig. 3. Counterweights are used to make the mechanism operation more efficient. Then the graph of the torque on the engine shaft against the angle of rotation changes and becomes similar to that given by curve 2 in Fig. 3. According to the methods developed above, the optimal (from the viewpoint of minimum energy losses in the engine) operation mode is defined by M = const, 2\u03c0M = mgh, (6.8) where m is the mass of oil lifted per one cycle and h is the hole depth. If the torque on the engine shaft is constant (curve 3 in Fig. 3), then an estimate of the losses obtained according to (5.6) and (5.7) shows that the optimal operation mode decreases the losses by 32% depending on the parameters of the conventional pumping unit. Moreover, such an actuator allows one to use a less powerful engine, because the constant torque on the shaft of an optimally operating engine is by 46% smaller than the maximum torque in the actuator operating under the same conditions. The above analysis and examples show that it is possible to pose the problem of constructing energysaving cyclic mechanisms" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003372_acs.macromol.0c02737-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003372_acs.macromol.0c02737-Figure1-1.png", "caption": "Figure 1. Chemical structures of bifunctional precursors and multifunctional cross-linkers for (a) PTHF and (b) PDMS gels.", "texts": [ ") and isocyanuric acid tris(5-pentyl isocyanate) as a bifunctional precursor and trifunctional cross-linker, respectively. PDMS gels were made using vinylterminated PDMS (Mn = 13,500 and Mw/Mn = 1.7; Gelest, Inc.) and tetrakis(dimethylsiloxy)silane as a bifunctional precursor and tetrafunctional cross-linker, respectively. For PDMS gels, a platinum catalyst (platinum(0)-1,3-divinyl-1,1,3,3-tetramethyldisiloxane) was used for the hydrosilylation reaction. The chemical structures of the precursors and cross-linkers are shown in Figure 1. The molar ratio of the functional group in the cross-linker (C) to that in the precursor chain (B) (r = [C]/[B]) in the reactant mixtures was varied in a range of r \u2264 1 to obtain specimens with various degrees of network sparsity. The values of r < 1 correspond to the conditions of excessive amounts of precursor, whereas r = 1 is the stoichiometric ratio. The sample codes G-X and Z-X in Table 1 represent the PTHF and PDMS gels with r values of X = 100 \u00d7 r, respectively. The reactant mixtures were stirred vigorously and subsequently degassed under vacuum" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001875_978-4-431-54595-8-Figure1.14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001875_978-4-431-54595-8-Figure1.14-1.png", "caption": "Fig. 1.14 Joint coordination driven by knee extensor #3 (Hosoda et al. 2010): (a) when extensor #3 contracts, (b) the knee is extended, (c) biarticular muscles #6 and #9 transfer force to hip and ankle joints, respectively, and (d) the whole leg is extended (This figure and caption are modified from Hosoda et al. (2010), with kind permission from Springer Science and Business Media)", "texts": [ " If the distal part is heavy, the eigenfrequency becomes low, which is a large disadvantage for fast motion. 20 K. Hosoda Biarticular muscle #9, gastrocnemius, drives knee and ankle joints. This muscle is also interesting from the mechanics viewpoint. It is not antagonistic to any muscle. We can conjecture that it is an evolutionally obtained body structure advantageous for forward moving, together with the knee locking mechanism. But the exact mechanics have not yet been fully revealed. Let us look at the joint coordination driven by knee extensor #3 (Fig. 1.14). When it contracts, the knee is extended (Fig. 1.14b). The extensor pulls biarticular muscle #6, and it extends the waist joint. The extensor also pulls biarticular muscle #9, and it extends the ankle joint (Fig. 1.14c, d). As a result, simply by contracting the knee extensor, the whole body is totally extended; this is obviously beneficial for jumping. When the robot touches the ground after a flight phase, the ankle is flexed, and gastrocnemius (#9) is pulled, causing the knee to flex. The flexed knee pulls muscle #5 and it flexes the waist joint. In total, the whole body is synergistically and automatically flexed. In the cases of jumping up and touching down, all the joints are coordinated by the complex muscular\u2013skeletal system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001555_978-3-319-18944-4_10-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001555_978-3-319-18944-4_10-Figure2-1.png", "caption": "Fig. 2. The simulation scene of mobile agents", "texts": [ " (30) serves as a feedback linearizing control for Eq. (28). Substituting Eq. (30) into Eq. (28), we have ide ide 1 1 1 ide 2 1 1 ( ) - ( ) ( ) r r k C K s s k \u03c1 \u03c1 \u03d5 \u03d5 \u2212 = = \u2212 = \u2212 rs (32) In order to demonstrate the validity of the proposed formation control approach, simulations are designed based on the hybrid platform of Webots 7 and Matlab. The leader-followers formation of four agents is formed where followers R2, R3 and R4 follow the leader R1. The simulation scenario given in Webots 7 is shown in Fig. 2. In the simulation, two typical trajectories are designed for the leader. As stated, to make the system always weakly observable ( 3), the trajectories do not include straight and parallel movements for the leader and followers. In scenario 1 (as shown in Fig.3 (a)), the trajectory of the leader is a complete elliptical. In scenario 2 (as shown in Fig.4 (a)), the leader will move straightly (shortly) firstly and then turn right and left in circles. Simulation results are given in subfigures of Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.19-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.19-1.png", "caption": "FIGURE 2.19 Tire under cornering conditions.", "texts": [ " The factor Dx is related to the peak of the longitudinal force coefficient (normalized longitudinal force) and the wheel load: Dx 5\u03bcxp Fz \u00f02:26\u00de Assuming pure longitudinal slip (no camber, no slip angle), this parame- ter \u03bcxp can be expressed in terms of Fz, as follows: \u03bcxp 5 \u00f0PDx1 1PDx2 dfz\u00de \u00f02:27\u00de for PDx1 and PDx2. Other parameters in the Magic Formula for pure longitudinal slip can be expressed as follows: Cx 5PCx1 \u00f02:28\u00de Bx Cx Dx 5Fz \u00f0PKx1 1PKx2 dfz\u00de exp\u00f0PKx3 dfz\u00de \u00f02:29\u00de Ex 5 \u00f0PEx1 1PEx2 dfz 1PEx3 df 2z \u00de \u00f012PEx4 sign\u00f0\u03ba\u00de\u00de \u00f02:30\u00de SHx 5PHx1 1PHx2 dfz \u00f02:31\u00de SVx 5PVx1 1PVx2 dfz \u00f02:32\u00de TABLE 2.3 Typical Values for the Nominal Tire Load Fz0 Class Fz0 [N] Compact class 3000 Middle class 5000 Top class 6000 Let us consider a tire under cornering conditions, as indicated in Figure 2.19 (top view), first neglecting camber. Under cornering conditions, a local velocity vector exists that is generally not parallel to the wheel center plane. This wheel center plane is defined as the symmetry plane of the tire such that forces acting in the symmetry plane do not contribute to a lateral force for the tire. In the front part of the contact area, the treads of the tire try to follow this local speed direction, resulting in a displacement from the symmetry plane along the tire circumference within the contact area, which increases linearly from zero (just in front of the contact area) up to a situation where the induced lateral shear stress just reaches the maximum possible shear stress level, i", " With different roll stiffnesses at front and rear axles, this works out differently at both axles, changing the handling characteristics of the vehicle. Lupton and Williams [22] give data for the sliding side force coefficient \u03bcys for one specific tire but different texture depths, under wetted conditions, and for two different speeds: 50 and 80 [km/h]. Results are shown in Figure 2.23. One observes some variation in results and dependency on texture depth. Also, observe the effect of speed: increased speed lowers the friction, especially with a small texture depth (as expected). Figure 2.19 indicates that the side force acts a small distance behind the wheel center. This distance is called the pneumatic trail tp(\u03b1). At small slip (small \u03b1), there is almost no sliding and the adhesion part of the contact area (linearly increasing lateral deflection) extends almost over the entire contact area. This corresponds to a situation where the shear stress profile is very asymmetrical along the contact area, with a rather large pneumatic trail. With increasing slip, the sliding area increases toward the front end of the contact area" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001620_s12206-015-0908-1-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001620_s12206-015-0908-1-Figure5-1.png", "caption": "Fig. 5. Static stress and strain state under low frequency cycle.", "texts": [], "surrounding_texts": [ "During the operation of an aero-engine, the turbine disc is subjected to a mixed load: centrifugal forces, thermal stresses, aerodynamic forces and vibratory stresses. Of course, high speed results in large centrifugal forces and high thermal gradients result in thermal stresses. Among them, the aerodynamic forces and vibratory stresses have little effect on the static strength of the turbine disc. Therefore, when analyzing the turbine disc with finite element method, the centrifugal forces and thermal stresses are the main consideration. The speed spectrum of the turbine disc is determined by the flight mission, and it consists of three parts [11]: low frequency cycle, full throttle cycle and cruise cycle. Any speed spectrum can be considered as a combination of these three basic cycles. The speed spectrum of the turbine disc is shown in Table 3. The temperature spectrum is derived based on the measurement data. In this study, the temperature spectrum of the turbine disc was loaded on the three-dimensional model by ANSYS parametric design language. For each basic cycle mentioned above, there are 100000 temperature data points of the turbine disc. Table 4 shows part of the temperature data points under full throttle cycle, where X, Y and Z represent the coordinate value of a point of the three-dimensional model." ] }, { "image_filename": "designv11_22_0001015_chicc.2014.6896862-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001015_chicc.2014.6896862-Figure1-1.png", "caption": "Fig. 1: The schematic diagram of a two-link manipulator.", "texts": [ " For all \u03c3 \u2265 1 and \u03c9 \u2208 R, the matrix (A\u2212 (\u03c3+ j\u03c9)BBTP is a Hurwitz matrix, where P is the symmetric positive de nite solution to the algebraic Riccati equation (14). Combining Theorem 3 and Lemma 2 leads to the following Theorem. Theorem 4. If G has a spanning tree, then by letting K = max{1, \u03c3\u22121 min}BTP where \u03c3min = min{ (\u03bb), \u03bb is the eigenvalue of L2} > 0, under the PI m D tracking protocol (7), all followers can asymptotically track the leader\u2019s trajecotry described by (8). In this section, a networked two-link manipulator system is considered. Four manipulator are treated as followers. The schematic diagram of the two-link manipulator is shown in Fig. 1, where mi, li are length and mass of linki, lci is the distance between the joint and the center of mass of linki, \u03b8i is the joint angle of linki. The dynamics of this two-link manipulator can be described by the following EL equation [20] M(q)q\u0308 + C(q, q\u0307)q\u0307 +G(q) = \u03c4, (15) where q(t) = (q1(t), q2(t)) T (\u03b81(t), \u03b82(t)) T , M(q) = [ M11(q) M12(q) M21(q) M22(q) ] , C(q, q\u0307) = [\u2212hq\u03072 \u2212hq\u03071 \u2212 hq\u03072 hq\u03071 0 ] , M11(q) =m1l 2 c1 + I1 +m2(l 2 1 + l2c2 + 2l1lc2 cos(q2)) + I2, M12(q) =M21(q) = m2l1lc2 cos(q2) +m2l 2 c2 + I2, M22(q) =m2l 2 c2 + I2, h =m2l1lc2 sin(q2), G(q) =(g1(q), g2(q)) T , G1(q) =m1lc1g cos(q1) +m2g[lc2 cos(q1 + q2) + l1 cos(q1)] G2(q) =m2lc2g cos(q1 + q2); and g is the gravitational acceleration, \u03c4 \u2208 R 2 is the input torque, Ii is the moment of inertia of linki" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure8.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0002985_s12206-020-1117-0-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure10-1.png", "caption": "Fig. 10. Finite element model of internal gear pair with single contact point: (a) contact condition; (b) finite element analysis model.", "texts": [ " To carry out the analysis of tooth contact, we needed to consider the interaction and deformation conditions between the tooth surfaces, so it was necessary to fix the outer surface of the internal gear. Furthermore, perfect constraints were imposed on the nodes of the outer surface of the internal gear (UX = UY = UZ = 0). And constraints were imposed on the degrees of freedom of all nodes on the inner surface of the pinion along the directions of UX and UZ, i.e., radial and axial directions (UX = UZ = 0). The finite element model of internal gears with single tooth is shown in Fig. 10. We analyzed the contact conditions of tooth profiles based on the setting conditions. The input torque applied to the inner surface of pinion was 200 Nm. Loading time was 1s and the time sub steps were set to 5. Stress analysis results of the internal gears with single contact point are displayed in Figs. 11-13. From the analysis results in Fig. 11, the maximum contact stress between the internal gear pair is 1429.4 MPa. The maximum stress occurs at contact point, which is located on the middle of the tooth profile" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000197_00450618.2019.1609088-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000197_00450618.2019.1609088-Figure2-1.png", "caption": "Figure 2. Typical components of a 3D printer.14", "texts": [ " The current range of consumer 3D printers has the potential to manufacture illicit objects thanks to the availability of 3D blueprints through a simple Google search.10 3D-printed illicit materials have been encountered in Australia with recent seizures of 3D-printed firearms and firearm components.11 In an attempt to combat the issue, the New South Wales State Government has implemented legislation making the downloading of instruction files for the 3D-printing firearms and firearm components illegal.12 Most 3D printers have the same basic construction design (Figure 2). They consist of a frame which holds the components together and supports the movement of the print head. The print head mechanism consists of an extruder that feeds the print filament through to the \u2018hot end\u2019, where it is heated before being deposited onto the print bed in layers along the X, Y and Z axis. The 3D printing process is computer controlled using CAD software to generate the 3D design.13 The process of 3D printing is undertaken in three phases: (1) The modelling phase consists of a digital design generated using computer-aided design (CAD) software" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001981_ijmpt.2016.079200-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001981_ijmpt.2016.079200-Figure2-1.png", "caption": "Figure 2 The main feature of a turning tool", "texts": [ " A considerable amount of investigations has been directed towards the prediction and measurement of cutting forces. That is because the cutting forces generated during metal cutting have a direct influence on the generation of heat, and thus tool wear, quality of machined surface and accuracy of the workpiece. Due to the complex tool configurations/cutting conditions of metal cutting operations and some unknown factors and stresses, theoretical cutting force calculations failed to produce accurate results. Figure 2 shows some of the main feature of a turning tool. The surface of the tool over which the chip flows is known as the rake face. The cutting edge is formed by the intersection of the rake face with the clearance face or flank of the tool. The tool is so designed and held in such a position that the clearance face does not rub against the freshly cut metal surface. The clearance angle is variable but is often of the order of 6\u00b0\u201310\u00b0. The rake face is inclined at an angle to the axis of the bar of work material and this angle can be adjusted to achieve optimum cutting performance for particular tool materials, work materials and cutting conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002701_aero47225.2020.9172614-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002701_aero47225.2020.9172614-Figure9-1.png", "caption": "Figure 9: Inertial and body-fixed frames of the Tetracopter.", "texts": [ " However, they still remain significantly under their critical buckling load, equal to 659N when computed with a length factor of 2. The Tetracopter is represented as a rigid body and its dynamics are derived with the Newton-Euler equations. The derivation of the nonlinear dynamic equations does not differ much from the derivations found in other papers that model flat quadcopters [8, 11, 3, 15]. However, the particular placement of the rotors implies a different expression of the torque induced by the differential thrust in (22). We consider an inertial reference frame and a body-fixed frame as shown on Fig. 9. The position and orientation of the body-fixed frame in the inertial frame is given by the translation vector \u03be and the Euler angles \u03b7 defined by (9) and (10). \u03be = [ x y z ] (9) Authorized licensed use limited to: Middlesex University. Downloaded on September 02,2020 at 09:13:47 UTC from IEEE Xplore. Restrictions apply. (b) Three-point bottom attachment scenario The rotation matrix from the inertial frame to the bodyfixed frame is given by (11), where the sines and cosines are abbreviated. R = [ c\u03c8c\u03b8 c\u03c8s\u03b8s\u03c6 \u2212 s\u03c8c\u03c6 c\u03c8s\u03b8c\u03c6 + s\u03c8s\u03c6 s\u03c8c\u03b8 s\u03c8s\u03b8s\u03c6 + c\u03c8c\u03c6 s\u03c8s\u03b8c\u03c6 \u2212 c\u03c8s\u03c6 \u2212s\u03b8 c\u03b8s\u03c6 c\u03b8c\u03c6 ] (11) The body-fixed frame has linear velocity V B = [u, v, w] and angular velocity \u03a9 = [p, q, r]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002599_s0263574720000685-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002599_s0263574720000685-Figure3-1.png", "caption": "Fig. 3. Cart-pole system.", "texts": [ " However, according to (40), we see that the derivative of the Lyapunov function of the closed loop system is negative semi-definite and d\u0302i (t)\u2212 di is bounded, which indicates that the update laws d\u0302i (t) finally converge to certain constant values. Here, in this section, we provide examples of the proposed motion planning and control algorithms for two PDF systems: the cart-pole system and the planar ballbot system. We choose a fourth-order Beta function in the motion planning examples and a set of polynomial basis functions for estimating the variation d in the control examples, that is, \u03c8i (x, t)= t i . The cart-pole system, as shown in Fig. 3, is a well-known test bed for nonlinear control techniques. It is combined by a sliding cart, on which a inverted pendulum is mounted, connected by a passive pin joint. The displacement of the cart is denoted by s and the pendulum angle by \u03b8 , increasing counterclockwise. The input force f is acting on the cart. The dynamic model of the cart-pole system is obtained1 as (m + m p)s\u0308 \u2212 m pl cos \u03b8 \u03b8\u0308 + m pl sin \u03b8 \u03b8\u03072 = f, (43) \u2212m pl cos \u03b8 s\u0308 + Jp \u03b8\u0308 + m pgl sin \u03b8 = 0, (44) where m, m p stand for the mass of the cart and of the pendulum and g for the standard gravitational acceleration" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002806_icra40945.2020.9197165-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002806_icra40945.2020.9197165-Figure3-1.png", "caption": "Fig 3: Sideslip-Yaw Rate Dynamic Model.", "texts": [ " The longitudinal error state (8), drops out and the lateral error state (9) becomes (11) when \ud835\udc65 is eliminated. (10) is restated as (12).\u220e Defining velocity heading error, ?\u0305? \ud835\udf03 \ud835\udefc , such that (11) becomes, \ud835\udc66 \ud835\udc63 sin ?\u0305? (13) which includes the effect of dynamics, \ud835\udefc , in ?\u0305? . Proposition 1: As in Fig 2, along the wheelbase, \ud835\udc63 cos \ud835\udefc \ud835\udc63cos \ud835\udefd . Considering small angles, \ud835\udc63 \ud835\udc63 . Therefore, the speed approximation \ud835\udc63 \ud835\udc63 is made. \u220e As in [15-17], vehicle dynamics are approximated by a bicycle model with cornering stiffness, Fig 3, which describes sideslip and yaw rate dynamics in the lateral direction [18]: \ud835\udefd \ud835\udc4e \ud835\udefd \ud835\udc4e \ud835\udc5f \ud835\udc4f \ud835\udf11 (14) \ud835\udc5f \ud835\udc4e \ud835\udefd \ud835\udc4e \ud835\udc5f \ud835\udc4f \ud835\udf11 (15) \ud835\udf11 \ud835\udf14 (16) where \ud835\udc4e , \ud835\udc4e 1 , \ud835\udc4e , \ud835\udc4e , \ud835\udc4f , \ud835\udc4f . The distances between the \ud835\udc36\ud835\udc3a and the front and rear axles is \ud835\udc3f and \ud835\udc3f , where \ud835\udc3f \ud835\udc3f \ud835\udc3f . Vehicle mass is \ud835\udc5a and inertia is \ud835\udc3d. Longitudinal tire forces are assumed to be small and are neglected. Lateral tire forces are approximated by \ud835\udc39 \ud835\udf07 \ud835\udc36 \ud835\udefc \ud835\udc36 \ud835\udefc , \ud835\udc39 \ud835\udf07 \ud835\udc36 \ud835\udefc \ud835\udc36 \ud835\udefc , where \ud835\udf07 and \ud835\udf07 are front and rear tire-ground friction coefficients. \ud835\udc36 and \ud835\udc36 are cornering stiffness" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002936_s12555-019-1073-6-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002936_s12555-019-1073-6-Figure1-1.png", "caption": "Fig. 1. Geometry of missile-target engagement.", "texts": [ " The 2D kinematics equation to represent an ASM model [10] is given by X\u0307m, j =Vm, j cos\u03d5m, j, (2) Y\u0307m, j =Vm, j sin\u03d5m, j, (3) where j is the index of the missile, Xm, j and Ym, j are the positions of the missiles in the inertial frame. Vm, j and \u03d5m, j denote the velocity and heading angle of the ASM, respectively. This study concentrates on a missile\u2019s terminal guidance phase; thus, it is assumed that its target information is known to the missile, guidance law is applied based on a missile engagement model. Fig. 1 shows the geometry of the missile engagement model. Proportional navigation guidance (PNG) is applied that as [16] ac, j = N \u2032Vc, j\u03bb\u0307 j, (4) where ac, j is an acceleration command, N \u2032 is a constant design factor that is usually determined between three and five, Vc, j denotes a missile-target closing velocity, and \u03bb j is a line-of-sight (LOS) angle. The closing velocity is defined as Vc, j = R\u0307s, j =Vs\u2212Vm, j, (5) where Rs, j is the relative distance between the jth missile and the target ship. The LOS angle can be calculated as \u03bb j = tan\u22121 ( Ys\u2212Ym, j Xs\u2212Xm, j ) " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000043_s12541-019-00047-7-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000043_s12541-019-00047-7-Figure11-1.png", "caption": "Fig. 11 The dynamic model of meshing-in impact", "texts": [ " After obtaining the deformation, we also need to determine the normal force that the current pairing teeth sustain at point M in 893 meshing periods. We can calculate the normal force Fni of point M by Eq.\u00a0(6). (5)\u0394xi = \u0394 i 3600 \u22c5 180 \u22c5 rbg \u22c5 cos b where hi and h\u02b9i are the load distribution factor between teeth at point M obtained by LTCA considering SRNPE when Tg= 500\u00a0N\u00a0m and T\u02b9g= 1500\u00a0N\u00a0m. The ratio from the normal force Fni to the deformation \u0394xi is the stiffness Ksi of point M, just like Eq.\u00a0(7), the value of which is shown in Fig.\u00a010. Figure\u00a011 shows the dynamic model of the meshing-in impact. We simplify the pinion and gear as particles. Their induced mass is, respectively: where Ip and Ig are the moment of inertia of the pinion and the gear. (6)Fni = T \ufffd g \u22c5 h\ufffd i \u2212 Tg \u22c5 hi rbg \u22c5 cos b (7)Ksi = Fni \u0394xi (8)mep = Ip r2 bp , meg = Ig r2 bg (9)Eki = 1 2 mev 2 si , (i = 1,\u2026 , 893) (10)Epi = \u222b Wsi 0 Ksixdx = 1 2 KsiW 2 si , (i = 1,\u2026 , 893) 1 3 where Eki is the impact kinetic energy; Epi is the elastic potential energy; vsi is the impact speed; me is the equivalent mass of the gear pair, a value can be obtained from Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001220_j.triboint.2014.10.021-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001220_j.triboint.2014.10.021-Figure4-1.png", "caption": "Fig. 4. Instrumented end pivoted roller follower.", "texts": [ "5 mm away and this posed a challenge keeping in mind that the roller by design floated in the housing axially due to the clearance of about 0.3 mm between the roller and the housing. Moreover, the relative high cost of the sensor, the complexity of installing it in the engine head and the durability issue due to high impact acceleration during valve operation limited its use in monitoring roller rotation speed in engine head. In the second technique, a miniature GMR sensor (Magnetometer) and a small magnet was used to monitor the roller rotation as shown in Fig. 4. Magnetometer sensors are very economical, easy to install and there is no adverse effect of high oil temperature and impact force on the performance of the sensor. The ultra-miniature size of the GMR sensor (1.1 mm 1.1 mm 0.45 mm) makes it the most suitable option for such space critical application. The third technique based on 5 mm Reed switch was also developed. In this method a miniature high response mechanical contact was used to monitor the roller speed. In the presence of a magnetic field, a contact is made in the miniature switch and when the magnetic field is removed the contact breaks", " Considering the performance of sensors and space available for mounting of sensor in roller follower assembly, GMR (magnetometer) sensor was selected for this research work. To convert the GMR chip into a sensor, a specially designed and fabricated printed circuit board (PCB) having 3 mm diameter was used for mounting the chip. A precision hole of 3.5 mm diameter was made in the roller housing using Computerized Numerical Control machine for mounting of sensor in such a manner that the hole was perfectly aligned with the roller race as shown in Fig. 4. For triggering of sensor, a very small Alnico magnet of 2 mm diameter and 0.5 mm thickness was inserted in the roller race. A thin layer of epoxy was applied for proper seating of magnet in the roller race. Special attention was paid to the orientation of the sensor in the housing and the clearance between the sensor and the target to avoid any kind of disturbance in the output signal. An advanced Data Acquisition system (DAQ) based on National Instruments (NI) hardware was used to monitor the roller rotation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002373_j.triboint.2020.106343-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002373_j.triboint.2020.106343-Figure2-1.png", "caption": "Fig. 2. Double slope bearing design with bent shaft.", "texts": [ " Misalignment values are often not zero, due to improper shaft alignment, excessive loading or other operational parameters. Misalignment angle can be resolved into two perpendicular angles, similar to the external load angles, one about each axis of the coordinate system according to Ref. [5]. Thus, lateral misalignment angles (\u03a8y) describe shaft rotations about the vertical y axis, and vertical misalignment angles (\u03a8x) describe rotations about the horizontal z axis as presented in our bearing model presented in Fig. 2. Eq. (3) will then be modified to: hhydro\u00f0\u03b8; z\u00de\u00bc c\u00fe e cos\u00f0\u03b8\u00de \u00fe z \ufffd \u03c8x cos\u00f0\u03b8\u00fe\u03d5\u00de\u00fe\u03c8y sin\u00f0\u03b8\u00fe\u03d5\u00de \ufffd (4) Bearing-shaft misalignment has a negative impact on the lubrication characteristics of the bearing. For a given pair of eccentricity and attitude angle values, the misaligned shaft is brought closer to the bearing surface, minimum film thickness decreases and the hydrodynamic lubrication film becomes less able to support additional radial loads. Another consequence of bearing-shaft misalignment is the alteration of the pressure distribution of the lubricant", " Tribology International 150 (2020) 106343 bearing film thickness, and will be used in the analysis of the bearing. Double slope design is proposed for stern tube bearings in order to maximize the contact area between the shaft and the bearing bush, resulting in decreased pressure through the length of the bearing. In comparison to other bearings of the shafting system, the aft stern tube bearing, due to the large L/D ratio, is more likely to require double slope instead of single slope inclination. The main reason for this, is that the shaft within the bearing will be elastically bent, as shown in Fig. 2. The elastic line of the bent shaft and the loads acting on the shaft are calculated utilizing an in house (NTUA) beam element based software for the calculation of the loads [4]. In order to model the elastically bent shaft, the shaft\u2019s length is divided into several segments and the misalignment parameter \u03a8x in Eq. (5) is replaced with a matrix of the respective inclination (\u03a8x values) of each longitudinal segment, calculated by the shaft alignment tool. The eccentricity value is calculated at the equivalent contact point location, taking into account local slope and misalignment, thus in single point location", "2 when: f ! i\u00f0 x!1\u00de\ufffd f ! i\u00f0 x!2\u00de 8i\u00bc 1;\u2026; k and f ! i\u00f0 x!1\u00de< f ! i\u00f0 x!2\u00de 9i\u00bc 1;\u2026; k The collection of all Pareto Dominant Solutions is called \u201cPareto Front\u201d. In this work, the design of double slope bearing is examined. The optimization aims at maximizing lubricant film thickness and minimizing the maximum pressure, with the first being equivalent to minimizing dimensionless eccentricity ratio. The optimization process includes the following three (3) dimensionless design parameters depicted in Fig. 2: 1. Secondary_slope_angle is the non-dimensional slope at the aft section of the bearing 2. Primary_slope_angle is the non-dimensional slope at the fore section of the bearing 3. Knuckle_point_position is the non-dimensional position of the knuckle point along the length of the bearing All parameters are dimensionless, since the slope angles are defined as a percentage of 2\u22c5c/L and the knuckle point position is defined as a percentage of the bearing length (%L). A case study will be presented, utilizing the tools introduced to study the stern tube bearing operation and design", " For a predefined loading condition and a shaft misalignment angle, the designs of no slope, single slope and double slope bearing will be compared for the linear and the bent shaft model. Then, single and double slope designs with bent shaft will be subjected to robustness tests, to demonstrate their performance in \u201cworse\u201d or G.N. Rossopoulos et al. Tribology International 150 (2020) 106343 overloaded conditions. In this case study, instead of conducting full shaft alignment calculations, a reduced model of the isolated stern tube bearing part of the shaft, with proper boundary conditions (as presented in Fig. 2), will be used for the calculation of the elastic line of the bent shaft in static loading conditions. All the initial parameters used in this case study are presented in Table 1. The purpose of this case study is to compare the bearing performance analysis results for several model types, with the goal of demonstrating the effect of single and double slope geometry on bearing performance. The process to evaluate the journal bearing performance parameters for both linear and bent shaft model is also included" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002725_tmag.2020.3021644-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002725_tmag.2020.3021644-Figure1-1.png", "caption": "Fig. 1. Existing CMGs. (a) CMG I, (b) CMG II.", "texts": [ " In our previous work, we found that a dual-flux modulated CMG (DFM-CMG) with radially magnetized PMs can also be realized [14]. The DFM-CMG can increase the torque capability. The purpose of this paper is to reveal the torque transmission mechanism in DFM configuration. The structure of the DFM-CMG is introduced in Section II. The operational principle and torque transmission are discussed in Section III. In Section IV, the torque characteristics are analyzed. In Section V, a DFM-CMG prototype is fabricated and tested. Finally, a conclusion is drawn in Section VI. Fig. 1 shows the two existing CMGs. The CMG I in Fig. 1(a) is a conventional configuration. Its modulator is sandwiched between the two rotors [1]. The CMG II in Fig. 1(b) was introduced in [15]. The modulator is adjacent to the inner side of LSR. Structurally, each of the CMGs meets [1, 15] LHm PPZ (1) where PH is the PPN of HSR PMs, PL is the PPN of LSR PMs, and Zm is the number of modulator segments. The air-gaps are marked in Fig. 1. The magnetic field in airgap II of CMG I contains the PH th and PL th order harmonics. If the HSR of CMG II is replaced with the HSR and modulator of CMG I, the PH th order harmonic in air-gap II can work as a HSR for CMG II, and the PL th order harmonic in air-gap II can couple with LSR PMs of CMG II. Hence, the DFM-CMG shown in Fig. 2 can be regarded as a combination of CMGs I and II. Due to the integration of two CMGs, the DFM-CMG structure is more complex. It consists of a HSR, an outer modulator (OM), a LSR, and an inner modulator (IM)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003741_j.promfg.2021.06.042-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003741_j.promfg.2021.06.042-Figure12-1.png", "caption": "Fig. 12. Theorized effects on the CA in the case of: (a) Over- (solid red) and under-compensated (dashed blue) beam offset (b) Over- (solid red) and under-scaled (dashed blue) X & Y beam positioning. (c) Superimposed effects of an over-compensated beam offset and under-scaled X & Y beam positioning. (d) Size error for case c. The dashed line represents the trend for size error if there was zero beam offset compensation error.", "texts": [ " Of course, while a direct machine evaluation approach might provide strong insight into beam positioning errors, the artifact evaluation approach allows for an analysis that can enable tuning of an AM system to achieve desired geometric characteristics such as size and position in an end-component. Examining Fig. 4a and Fig. 5a, which show the LA and CA size errors, should illuminate beam offset errors. Barring the influence of other factors, a consistent beam offset error should result in a consistent feature size error. These errors should be opposite in sign depending on whether the feature is positive, such as a boss, or negative, such as a bore (see Fig. 12a). That said, if significant beam positioning error effects are present, this prevents a simple application of this rule for all features of size throughout the build volume. As mentioned earlier, considering the bias between LA feature size errors (Fig. 4a) relative to feature position errors (Fig. 4b) is an approximation for removing beam positioning error effects. In this case the negative bias suggests an over-compensated beam offset. An important caveat to this diagnosis is that the small feature size and sparse surface sampling of the LA could have very well also contributed to exhibited biases \u2013 the relatively noisy trends in all LA measurements exemplify this", " Taken together, CA1, CA2, and C-RP errors all indicate gross beam positioning scale error, as larger features generally have more negative size error. The +XLA measurements in Fig. 10b support this claim, as they show that the AM system is undershooting its commanded beam position. Additionally, the CA2 features (located further from the origin) are slightly more undersized than the CA1 features, which follows with the +X scaling error suggested by Fig. 4b. Consider the simplified cases presented in Fig. 12. Several simplifying assumptions are made \u2013 X and Y beam-positioning error is assumed to scale linearly with distance from the build area origin and the CA is considered to be centered on the origin. Fig. 12a displays the expected changes should an overor under-compensated beam offset be present \u2013 both rings either shrink or grow due to a uniform translation of all feature surfaces. Fig. 12b shows the results of over-scaled or underscaled X and Y beam positioning, resulting in all surfaces moving out or in with error magnitude scaling with feature size. Fig. 12c provides an example of the superimposed effects of an over-compensated beam offset and under-scaled beam positioning error \u2013 a condition that appears to be present in the examined AM machine. Plotting these diameter errors would result in the plot shown in Fig. 12d, assuming no effects of random process fluctuations. Comparing to Fig. 5a confirms that the beam positioning system is undershooting its commanded value, but either a minimal beam offset error is present (thus the pattern in Fig. 12d is not shown) or other confounding effects dominate. The latter seems most likely, due a lack of consistency in how the boss-like and bore-like features are related to each other in Fig 5a. Those applying this standard should use caution if they look to the smallest positive resolution-type features to study beam offset error. In this case local temperature rise during the AM process may significantly affect meltpool geometry. Negative features, such as the RH and RS, should also be disregarded when assessing beam offset error" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003429_j.tws.2021.107585-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003429_j.tws.2021.107585-Figure3-1.png", "caption": "Fig. 3. Geometry description and loads of an arbitrary shell structure.", "texts": [ " The internal forces in the form of forces and moments are following: \ud835\udc411 = \ud835\udc361 ( \ud835\udf001 + \ud835\udf082\ud835\udf002 ) , \ud835\udc412 = \ud835\udc362 ( \ud835\udf002 + \ud835\udf081\ud835\udf001 ) , (21) \ud835\udc401 = \ud835\udc371 ( \ud835\udf121 + \ud835\udf082\ud835\udf122 ) , \ud835\udc402 = \ud835\udc372 ( \ud835\udf122 + \ud835\udf081\ud835\udf121 ) , (22) where: \ud835\udc361 = \ud835\udc381\ud835\udc61 1 \u2212 \ud835\udf081\ud835\udf082 , \ud835\udc362 = \ud835\udc382\ud835\udc61 1 \u2212 \ud835\udf081\ud835\udf082 , (23) \ud835\udc371 = \ud835\udc381\ud835\udc613 12(1 \u2212 \ud835\udf081\ud835\udf082) , \ud835\udc372 = \ud835\udc382\ud835\udc613 12(1 \u2212 \ud835\udf081\ud835\udf082) , (24) n which: \ud835\udc381, \ud835\udc382 \u2014 Young\u2019s modulus, \ud835\udf081, \ud835\udf082 \u2014 Poisson\u2019s ratio. The rincipal stresses are defined as follows: 1 = \ud835\udc361 \ud835\udc61 ( \ud835\udf001 + \ud835\udf082\ud835\udf002 ) , \ud835\udf0e2 = \ud835\udc362 \ud835\udc61 ( \ud835\udf002 + \ud835\udf081\ud835\udf001 ) . (25) tress across the thickness of a shell: 1\ud835\udc67 = \ud835\udc361 \ud835\udc61 ( \ud835\udf001\ud835\udc67 + \ud835\udf082\ud835\udf002\ud835\udc67 ) , \ud835\udf0e2\ud835\udc67 = \ud835\udc362 \ud835\udc61 ( \ud835\udf002\ud835\udc67 + \ud835\udf081\ud835\udf001\ud835\udc67 ) . (26) Strain energy is formulated in the following manner \ud835\udc48\ud835\udf00 = \ud835\udf0b \u222b \ud835\udf112 \ud835\udf111 ( \ud835\udc411\ud835\udf001 +\ud835\udc412\ud835\udf002 +\ud835\udc401\ud835\udf121 +\ud835\udc402\ud835\udf122 ) \ud835\udc5f d\ud835\udc5f d\ud835\udf11 1 cos\ud835\udf11 d\ud835\udf11, (27) where: \ud835\udf111, \ud835\udf112 are the coordinates describing boundaries of a shell (Fig. 3). 3.2. Orthotropic cylindrical shells In the case of cylindrical shells, most of the formulae from the previous section can be significantly simplified. Strains definition can be rewritten to the form: \ud835\udf001 = 1 d\ud835\udc62 , \ud835\udf002 = \u2212 1 \ud835\udc64, (28) \ud835\udc3f d\ud835\udf09 \ud835\udc450 t \ud835\udf00 4 4 m s f \ud835\udef1 m t a s t f \ud835\udc4a I p c c s \ud835\udc41 { ( )\u22123 \u2212 Importantly, the vertical and horizontal displacements are coherent with normal and tangent directions correspondingly. The changes in curvature are following: \ud835\udf121 = 1 \ud835\udc3f2 d2\ud835\udc64 d\ud835\udf092 , \ud835\udf122 = 0, (29) herefore the strains across the thickness: 1\ud835\udc67 = \ud835\udf001 + \ud835\udc67\ud835\udf121, \ud835\udf002\ud835\udc67 = 1 1 \u2212 \ud835\udc67 \ud835\udc450 \ud835\udf002", " Strain energy of the cylindrical shell is formulated analogously to the doubly curved shells \ud835\udc48\ud835\udf00 = \ud835\udf0b\ud835\udc3f\u222b \ud835\udf09\ud835\udc502 \ud835\udf09\ud835\udc501 ( \ud835\udc411\ud835\udf001 +\ud835\udc412\ud835\udf002 +\ud835\udc401\ud835\udf121 +\ud835\udc402\ud835\udf122 ) \ud835\udc5fd\ud835\udf09. (31) . The Ritz method application .1. Theoretical description The Ritz method is described widely in the literature [1,2]. The ethod is based on minimization of the total potential energy of a ystem \ud835\udef1 expressed as strain energy \ud835\udc48\ud835\udf00 and potential energy of external orces \ud835\udc4a\ud835\udc5d = \ud835\udc48\ud835\udf00 \u2212\ud835\udc4a\ud835\udc5d. (32) An arbitrary shell structure of revolution is determined by the eridional angle in the range \ud835\udf111 \u2264 \ud835\udf11 \u2264 \ud835\udf112 (Fig. 3). The structure is loaded on its edges with normal forces \ud835\udc4b1, \ud835\udc4b2, ransverse loads \ud835\udc4c1, \ud835\udc4c2 and moments \ud835\udc4d1, \ud835\udc4d2, while uniform pressure \ud835\udc5d cts on the middle surface of the shell. To apply the Ritz method for uch a problem, it is necessary to describe the strain energy and potenial energy of the external forces as functions of unknown displacements unctions. The latter is expressed as \ud835\udc5d = 2\ud835\udf0b [ 2 \u2211 \ud835\udc56=1 \ud835\udc4b\ud835\udc56\ud835\udc5f ( \ud835\udf11\ud835\udc56 ) \ud835\udc62(\ud835\udf11\ud835\udc56) + 2 \u2211 \ud835\udc56=1 \ud835\udc4c\ud835\udc56\ud835\udc5f ( \ud835\udf11\ud835\udc56 ) \ud835\udc64(\ud835\udf11\ud835\udc56) + 2 \u2211 \ud835\udc56=1 \ud835\udc4d\ud835\udc56\ud835\udc5f ( \ud835\udf11\ud835\udc56 ) \ud835\udf171(\ud835\udf11\ud835\udc56) \u2212 \ud835\udc5d\u222b \ud835\udf112 \ud835\udf111 \ud835\udc64\ud835\udc5f d\ud835\udc5f d\ud835\udf11 1 cos\ud835\udf11 d\ud835\udf11 ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003684_j.matpr.2021.05.466-Figure28-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003684_j.matpr.2021.05.466-Figure28-1.png", "caption": "Fig. 28. Circumferential stress for stainless steel.", "texts": [], "surrounding_texts": [ "Figs. 16, 17, 18, and 19 shows the result of longitudinal stressinduced in Elliptical, Hemispherical, Torispherical, and Plain formed head. Fig. 23. Circumferential stress for low carbon steel. Fig. 24. Longitudinal Stress for alluminium alloy. Fig. 26. Circumferential stress for gray cast iron. Fig. 27. Longitudinal stress for gray cast iron. Fig. 29. Longitudinal stress for stainless steel. Fig. 30. Longitudinal stress for titanium alloy. Fig. 31. Circumferential stress for titanium alloy. From graph Fig. 20, Stress-induced in Hemispherical head is minimum, while in Plain formed head, stress-induced is maximum. Therefore, for the taken conditions, 8 bar capacity pressure vessels and 24 L, the hemispherical end is better." ] }, { "image_filename": "designv11_22_0000465_0954406219878755-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000465_0954406219878755-Figure1-1.png", "caption": "Figure 1. Enveloping conical worm pair.", "texts": [ " Besides, when the center distance of the worm pair is less, the transmission ratio is larger and the number of thread of the worm is more, the meshing limit line may be closer to the little end of the conical worm e helicoid. Enveloping conical worm pair, meshing limit line, characteristic point, nonlinear equation set, conical surface Date received: 17 March 2019; accepted: 25 August 2019 The conical surface enveloping conical worm pair, which can be named \u2018\u2018enveloping conical worm pair\u2019\u2019 in short and used for transmitting power and motion between two crossed shafts, is a new-style mechanical transmission device.1\u20133 As displayed in Figure 1, a conical worm pair includes an enveloping conical worm and a conical worm wheel. The helicoid of the conical worm is enveloped by a discoid conical grinding wheel according to its formation mechanism. The related worm wheel is formed by a conical hob, whose generating surface is identical to the preceding conical worm helicoid. For the enveloping conical worm pair, its meshing principal has been investigated in depth by Zhao and Kong.1 The axial tooth form and the computation of the angle of tooth form for the enveloping conical worm have been studied in literature" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003676_s00170-021-07413-8-Figure36-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003676_s00170-021-07413-8-Figure36-1.png", "caption": "Fig. 36 A sound formed St13 conical cups obtained from a simulation b experiment, Pcavity max=9 MPa and Pradial max=13 MPa, punch displacement= 24 mm", "texts": [ " According to the studies in the previous sections and finding the effect of radial and the cavity pressures on wrinkling, and by controlling the maximum wrinkling height parameter as a result of pressure variation, a piece without wrinkling and rupture was formed. For preventing the wrinkling on the wall, various simulations have been performed with different pressure paths in HDDRP with inward flowing liquid. The main difference between the pressure paths was maximum pressure. The comparison between the maximum wrinkling height data shows that for the cavity pressure of 9 MPa and the radial pressure of 13 MPa, no wrinkling happened, as shown in Fig. 36. However, forming a steel cone with a high LDR (2.55 in this study) while having a large wall area is not possible in conventional hydroforming processes [8, 18], but by using this process and its ability, the cone was formed without rupture, but still had a significant amount of wrinkling. Then, according to the results stated in the previous section, and by controlling the maximum wrinkling height parameter by changing the pressure, a piece without any wrinkle or rupture was formed, as shown in Fig. 36. In this study, the effects of the cavity and radial pressures on copper and St13 conical parts formed in the hydrodynamic deep drawing process assisted by radial pressure and inward flowing liquid were investigated. Wrinkling wave in the formed conical cups was studied using finite element simulation. To verify the FE simulation results, several experiments were performed. The results showed that by increasing the cavity pressure or decreasing the radial pressure, the maximum wrinkle height in the formed conical cups decreases" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002105_iros.2016.7759528-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002105_iros.2016.7759528-Figure2-1.png", "caption": "Figure 2. Diagram of simple kinematic differential drive", "texts": [ " \ud835\udef4\ud835\udc5d\ud835\udc52\ud835\udc60\ud835\udc61 is the covariance matrix of the parameter estimate and \u2206\ud835\udc58 is the covariance matrix that governs how much weight we put on the most recent measurement. The larger that \u2206\ud835\udc58 is, the more weight we put on the recent measurement. We can also estimate for \ud835\udef4\ud835\udc64, assuming that it is a constant, although it can vary according to terrain type. This will not be discussed in detail due to limited space, but more information can be found in [10]. In this paper, the following simple kinematic differential drive with augmented input is of interest. Figure 2 shows the diagram of the vehicle. [ ?\u0307? ?\u0307? ?\u0307? ] = [ (\ud835\udefd + \ud835\udc64\ud835\udefd) \ud835\udc63\ud835\udc5f+\ud835\udc63\ud835\udc59 2 \ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf03 (\ud835\udefd + \ud835\udc64\ud835\udefd) \ud835\udc63\ud835\udc5f+\ud835\udc63\ud835\udc59 2 \ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf03 (\ud835\udefc + \ud835\udc64\ud835\udefc) \ud835\udc63\ud835\udc5f\u2212\ud835\udc63\ud835\udc59 \ud835\udc4f ] = [ \ud835\udc63\ud835\udc5f,\ud835\udc4e+\ud835\udc63\ud835\udc59,\ud835\udc4e 2 \ud835\udc50\ud835\udc5c\ud835\udc60\ud835\udf03 \ud835\udc63\ud835\udc5f,\ud835\udc4e+\ud835\udc63\ud835\udc59,\ud835\udc4e 2 \ud835\udc60\ud835\udc56\ud835\udc5b\ud835\udf03 \ud835\udc63\ud835\udc5f,\ud835\udc4e\u2212\ud835\udc63\ud835\udc59,\ud835\udc4e \ud835\udc4f ] where \ud835\udc67 = [\ud835\udc65, \ud835\udc66, \ud835\udf03]\ud835\udc47 is the state vector, \ud835\udc62 = [\ud835\udc63\ud835\udc5f , \ud835\udc63\ud835\udc59] \ud835\udc47 is the nominal input vector applied to the system, and \ud835\udc62\ud835\udc5d = [\ud835\udc63\ud835\udc5f,\ud835\udc4e \ud835\udc63\ud835\udc59,\ud835\udc4e] \ud835\udc47 is the augmented input vector. \ud835\udc5d = [\ud835\udefc \ud835\udefd]\ud835\udc47 is the parameter vector and \ud835\udc64 = [\ud835\udc64\ud835\udefc \ud835\udc64\ud835\udefd]\ud835\udc47 is the process noise vector which takes a normal distribution, \ud835\udc64~\ud835\udc41(0, \ud835\udef4\ud835\udc64) where \ud835\udef4\ud835\udc64 is a constant matrix. \ud835\udc62 and \ud835\udc62\ud835\udc4e satisfy the following equation: \ud835\udc62\ud835\udc4e = [ (\ud835\udefc+\ud835\udc64\ud835\udefc)+(\ud835\udefd+\ud835\udc64\ud835\udefd) 2 \u2212 (\ud835\udefc+\ud835\udc64\ud835\udefc)\u2212(\ud835\udefd+\ud835\udc64\ud835\udefd) 2 \u2212 (\ud835\udefc+\ud835\udc64\ud835\udefc)\u2212(\ud835\udefd+\ud835\udc64\ud835\udefd) 2 (\ud835\udefc+\ud835\udc64\ud835\udefc)+(\ud835\udefd+\ud835\udc64\ud835\udefd) 2 ] \ud835\udc62 = \ud835\udc47\ud835\udc4e\ud835\udc62 The vehicle model can be considered as the effective wheel base model with an additional parameter \ud835\udefd" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure6.20-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure6.20-1.png", "caption": "Figure 6.20 Vernier protractor scale readings", "texts": [ " The moving scale is teeshaped\u00a0to provide a substantial base and datum from which readings are taken. The instrument reading is the amount which the rule sticks out beyond the base. Depth gauges are available in a range of capacities from 150 mm to 300 mm. Figure 6.13 Vernier height guage Figure 6.14 Digital height gauge D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 6 Measuring equipment 6 88 vernier scale is 23 12 degrees = 111 12 degrees or 1\u00a0degree 55 minutes. This is 5 minutes less than two divisions on the fixed scale (Fig. 6.20(a)). If the two scales initially have their zeros in line and the vernier scale is then moved so that its first graduation lines up with the 2 degree graduation on the fixed scale, the zero on the vernier scale will have moved 5 minutes (Fig. 6.20(b)). Likewise, the second graduation of the vernier lined up with the 4 degree graduation will result in the vernier scale zero moving 10 minutes (Fig. 6.20(c)) and so on until when the twelfth graduation lines up the zero will have moved 12 \u00d7 5 = 60 minutes = 1\u00a0degree. Since each division on the vernier scale represents 5 minutes, the sixth graduation is numbered to represent 30 minutes and the twelfth to represent 60 minutes. As well as linear measurement, vernier scales can equally well be used to determine angular measurement. The vernier bevel protractor (Fig.\u00a06.18) again uses the principle of two scales, one moving and one fixed. The fixed scale is graduated in degrees, every 10 degrees being numbered 0,10, 20, 30, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002888_s12221-020-1016-0-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002888_s12221-020-1016-0-Figure7-1.png", "caption": "Figure 7. Derivation of thickness (a) outside-in toolpath and (b) inside-out toolpath.", "texts": [ " The experiments continued normally until the plate was broken. In the three-tier-plate clamping method, the forming depth of the upper iron plate had a positive impact on the forming depth of composite sheet. Thus, it was necessary to investigate the effect of toolpath on forming depth of the iron upper plate. Cosine law model was commonly used in the thickness prediction of metal sheet under forming [20]. In the three-tier-plate clamping method, the thickness of the upper iron plate accorded with the cosine law model. As shown in Figure 7, based on volume invariably, the thickness of the iron plate under the two toolpaths could be obtained: Outside-in (1) Inside-out (2) where t0 represents the initial thickness, t1 denotes the final thickness of the side wall in outside-in toolpath, t2 denotes the final thickness of the side wall in inside-out toolpath, dx represents a small initial width of the initial plate, ds represents the final width corresponding to the small initial width, S0 represents the length of the bottom of pyramid, S1 represents half the length of the bottom of the pyramid" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001083_1.4027926-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001083_1.4027926-Figure9-1.png", "caption": "Fig. 9 Test-site photo of compressor rotor experimental rig", "texts": [ " These faults are produced on the inner ring, outer ring, and ball of the compressor experimental rig\u2019s 6214 ball bearing and the aero-engine experimental rig\u2019s 6205 ball bearing, which are shown in Figs.7 and 8. The ball bearing dimensions are listed in Table 1. The ball bearing faults\u2019 characteristics frequencies can be computed as follows: (1) Outer race foc \u00bc Z 2 1 d D cos a fR (1) (2) Inner race fic \u00bc Z 2 1\u00fe d D cos a fR (2) (3) Rolling ball fbc \u00bc Z 2 D d 1 d D cos a 2 \" # fR (3) (4) Cage fc \u00bc 1 2 1 d D cos a fR (4) 4.2 Faults Experiments of Ball Bearings. The bearing faults experiments are carried out. The test-site photo of the compression rotor rig is shown in Fig. 9, and the measurement points\u2019 explanation is listed in Table 2; the test-site photo of the aeroengine rotor rig is shown in Fig. 10, and the measurement points\u2019 explanation is listed in Table 3. Vibration signals are collected by means of the USB9234 data acquisition card of the NI Company, the 4805 type ICP acceleration sensors of B&K Company are used to pick up the acceleration signals, and the eddy current sensors are used to measure the rotating speeds. The sampling frequency is 10.24 kHz. 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002715_j.polymer.2020.122973-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002715_j.polymer.2020.122973-Figure14-1.png", "caption": "Fig. 14. True stress in the loading direction along the outer surface for: a) dumbbell, b) concave.", "texts": [ "11 higher than for the dumbbell test specimen. Fig. 13 shows the true stress in the direction of loading at a load of 200 N on both specimens. The stress distribution clearly shows that stress concentration occurs in the transition area of the dumbbell (grey areas). In the case of the concave, the stress concentration area is rather negligible. Both test specimens show identical stress values in the homogeneous deformed area. The true stress value in the homogeneous area is about 2 for both test specimens. Fig. 14 shows the true stress in the direction of loading as a function of the distance of the undeformed path along the specimen. The stress concentration strongly influences crack initiation and crack growth of the dumbbell test specimen. Most dumbbell test samples failed near the transition area, indicating that the stress concentration is the dominant factor in crack initiation. For the dumbbell, the stress concentration factor is about 1.29. For the concave, the stress increases by a factor of 1.05" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003717_s00521-021-06215-z-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003717_s00521-021-06215-z-Figure1-1.png", "caption": "Fig. 1 The tilting quadcopter vehicle", "texts": [ " Finally, we draw a conclusion in Sect. 5. Notations: Throughout this paper, let R denote the set of real numbers and Rn m denote a n m-dimensional matrix. Let Ki \u00bc diag\u00bdkij 2 R3 3 and Ki \u00bc diag\u00bd kij 2 R3 3 denote positive-definite matrices with i \u00bc 1; 2; ; 6; j \u00bc 1; 2; 3. Let zk \u00bc \u00bdjz1jksgn\u00f0z1\u00de; ; jznjksgn\u00f0zn\u00de T, where z \u00bc \u00bdz1; ; zn T and sgn\u00f0 \u00de is the sign function. In this section, we propose a novel conception of the tilting quadcopter, which consists of four main rigid parts: the Body, Servo, Push and Propeller as shown in Fig. 1. Servo and Push mechanisms can be actuated to tilt about the corresponding axes w.r.t. the main body, which suggests a fully actuated vehicle can be obtained. Compared with [18] and [19], the tilting mechanisms of this novel conception can be indeed arbitrary within \u00bd p 2 ; p 2 to provide more agile and reliable abilities. Let RW : fow; xw; yw; zwg and RB : fob; xb; yb; zbg denote the World reference frame and the Body frame. Let RSi : fosi ; xsi ; ysi ; zsig and RPi : fopi ; xpi ; ypi ; zpig; i \u00bc 1 4 denote the Servo and Push frames associated with the i-th propeller as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000902_ilt-01-2015-0003-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000902_ilt-01-2015-0003-Figure5-1.png", "caption": "Figure 5 Schematic of top compression ring and cylinder liner contact surface", "texts": [ " h0 = t h0 t h0 = t 10 5 (29) h0 = T h0 = 0 h0 = T 10 5 (30) Equation (9) indicates that the accuracy of the Archard\u2019s wear law mainly depends on the wear coefficient in the boundary lubrication regime k, which is related with the surface topography, the materials, the lubrication and the operating conditions (Priest et al., 1999). In this research, a reciprocating friction and wear test rig is used to obtain k during running-in process, while verifying the validation of the uneven wear model. The schematic of the piston ring and cylinder liner contact surface is shown in Figure 5. To ensure circumferential conformity, the larger diameter cylinder liner with whose topography is similar with the real one, is used in the experiment. In total, 20-mm wide and 30-mm long cylinder Figure 15 On the running-in behavior of rough surface Jun Cheng, Xianghui Meng, Youbai Xie and Wenxiang Li Industrial Lubrication and Tribology Volume 67 \u00b7 Number 5 \u00b7 2015 \u00b7 468\u2013485 liner specimen and 10-mm long ring specimen cutting from the top compression ring are fixed in the holders. The cylinder liner is grey cast iron and the piston ring is DLC Cr coated low alloy steel" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001170_0278364914551773-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001170_0278364914551773-Figure4-1.png", "caption": "Fig. 4. Four wheel steering. All four wheels are steerable. As a result, this vehicle can drive in any direction no matter where it is pointing. This figure suggests driving left while pointing forward. The wheel contact points are offset by a distance d from the steering centers.", "texts": [ " This solution produces the linear and angular velocities that are most consistent (in a least-squares sense) with the measurements, even if the measurements do not agree. This section provides some non-trivial worked examples of velocity kinematics including comparisons with other methods. This is the case of four independently steerable wheels. Subject to any limits on steering angles, this vehicle configuration is very maneuverable. It can turn in place and it can drive in any direction while facing another direction (Figure 4). A related configuration is double Ackerman. In that case, there is an Ackerman mechanism at both the front and the rear and such a vehicle can turn through smaller radii than single Ackerman with the same steering limits. For the four wheel steering configuration, let the wheel frames be identified by numbers as shown. The centers of their contact points are assumed to be in the center of their vertical projections. Let the forward velocity of the vehicle frame be denoted vx and the angular velocity is v", " The parameters used will depend on which of the six chains is being computed. Simulation results for Rocky7 are presented in the next section. In order to illustrate generality, we chose to verify the model of a very complicated four-wheel-steering vehicle executing a very complicated trajectory. Nonetheless, even this difficult case is fairly straightforward using our methodology. The second experiment provides numerical simulation results for the Rocky7 platform moving over rolling terrain. We simulated the four-wheel-steering vehicle depicted in Figure 4 following a hypotrochoid trajectory. This is the curve traced by a point attached to a small circle with radius r as it rolls along the inside of a larger circle with radius R. The equations for the (x,y) coordinates of the vehicle frame as a function of time are therefore x(t)= (R r) cos (t)+ r cos R r r t \u00f0103\u00de y(t)= (R r) sin (t) r sin R r r t \u00f0104\u00de For the trajectories in Figures 7 and 8, r = 8, R = 10, and r = 3.2. The overall simulation proceeds by first generating the ground truth linear and angular velocity from the above analytic curve" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001683_0954406216639075-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001683_0954406216639075-Figure2-1.png", "caption": "Figure 2. Determination of normal chordal dimension SFn of tooth root critical section and bending moment arm hFe. 2", "texts": [ "0 when 6\u00bc 0 and when > 30 , \u00bc 30 and when overlap ratio \" > 1, \" \u00bc 1. The extended helix factor Y 0 considers the proper tooth depth of helical gears. The factor is calculated according to the following equation. Y 0 \u00bc Y 1 cos3 \u00f04\u00de The form factor YF, which takes into account the influence on nominal tooth root stress of the tooth form with load applied at the outer point of single pair tooth contact. It is predicted by ISO 6336:3- 2006 in following equation. YF \u00bc 6hFe mn cos Fen SFn mn 2 cos n \u00f05\u00de The symbols used in above equation are illustrated in Figure 2. The extended form factor YF 0 is corrected by the additional factor f\". Y0F \u00bc YFf\" \u00f06\u00de The factor f\" considers the influence of load distribution between the teeth in the mesh. The factor is calculated according to the following equations. If \" \u00bc 0 and \" n< 2 then f\" \u00bc 1 \u00f07\u00de If \" \u00bc 0 and \" n5 2 then f\" \u00bc 0:7 \u00f08\u00de If 0<\" < 1 and \" n< 2 then f\" \u00bc 1 \" \u00fe \" \" n 0:5 \u00f09\u00de If 0<\" < 1 and \" n5 2 then f\" \u00bc 1 \" 2 \u00fe \" \" n 0:5 \u00f010\u00de If \" 5 1 then f\" \u00bc \" 0:5 n \u00f011\u00de The proposed method and the original method have been tested in the work reported in this paper" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002491_1077546320932030-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002491_1077546320932030-Figure4-1.png", "caption": "Figure 4. First six relative empirical mode decomposition energy entropies of different operating conditions.", "texts": [ " The vibration signals under those four conditions are selected as samples, and 100 bearings for each state were tested. Thus, 400 data can be obtained, and each containing 25,000 sample points. The training dataset is half samples of the original dataset in the experiment. After the high-dimensional feature set is constructed, it is input into CIKECA for eliminating redundant and extracting low-dimensional features, and the first six relative EMD energy entropies of different operating conditions are shown in Figure 4. To examine whether such an introduction of class label to KECA is helpful for recognition, similar methods including PCA, KPCA, and KECA are conducted for comparison. The target dimension for each method is set to a certain number so that the cumulative variance contribution rate is more than 95%. The Gaussian kernel function is selected for all the kernel functions, and the five-fold cross-validation is applied to decide parameter \u03c3, where the parameter \u03c3 of KPCA, KECA, and CIKECA is 1.2, 1.5, and 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000092_s12541-019-00032-0-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000092_s12541-019-00032-0-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of laser\u2013TIG hybrid welding [25]", "texts": [ " For the experiments in this study, 1.5\u00a0mm thick commercial AA6061-T651 aluminum alloy rolled plates were used. The chemical composition of AA6061-T651 is shown in Table\u00a01. The tensile strength of the Al alloy is 340\u00a0MPa. The welding process was produced by Nd:YAG laser\u2013induced TIG hybrid welding technique. Nd:YAG laser and TIG welding machine were used in experiment. The distance and angles between the laser-focused point and the TIG torch was fixed at 2\u00a0mm and 60\u00b0, respectively. A schematic drawing of the process is shown in Fig.\u00a01. As shown in Fig.\u00a01, the combined method of laser front is adopted. Under the same welding process parameters, when the laser front is used, the welding surface is bright, the fishscale grain is fine and smooth, the welding process is stable, and the welding seam is well formed. When the arc front is used, the surface of welding seam becomes black, oxidation is serious, and the forming of welding is unstable. The parameters of the pulse laser were as follows: 1.064\u00a0\u03bcm wavelength, 120\u00a0mm core diameter, 0\u00a0mm defocusing distance" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000075_asjc.2042-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000075_asjc.2042-Figure1-1.png", "caption": "FIGURE 1 Fully actuated hexarotor configuration [Color figure can be viewed at wileyonlinelibrary.com]", "texts": [ " Under these assumptions, the rotational dynamic model of the multirotor can be reduced to ?\u0308? = I\u22121 B \ud835\uded5. (10) The second-order dynamic models (9) and (10) are more appropriate for the control system design and for the algebraic estimator and observer design. In [17,18], a design and control of fully actuated passively tilted multirotor are proposed. The non-flat design with passively tilted rotors can overcome the inherent underactuated property of the flat multirotor configurations, as shown in Figure 1. The passively tilted multirotor is able to achieve full controllability and decoupling position from orientation. This fact has a significant influence on the multirotor controller design. The non-flat configuration provides six independent control variables, the one for each degree of freedom, contrary to the flat configuration, which provides only four independent control variables [19]. The feedback control law based on the fully actuated multirotor, which provides asymptotic tracking of desired trajectory xd(t) is F = mR(\ud835\udf3c)T[x\u0308d + ge3 \u2212 KD \u0307\u0303x \u2212 KPx\u0303], (11) so that the closed-loop tracking error x\u0303 = x \u2212 xd satisfies \u0308\u0303x + KD \u0307\u0303x + KPx\u0303 = 0, (12) which is asymptotically stable for the positive-definite gain matrices KD,KP \u2208 R3\u00d73 [20]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002242_j.ijsolstr.2020.01.021-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002242_j.ijsolstr.2020.01.021-Figure2-1.png", "caption": "Fig. 2. A four-stage tensegrity tower.", "texts": [ " (4) Under the driving of the selected active members, the deviation between the final configuration and the target configuration can be estimated accurately. However, the active members selected using the proposed method cannot be guaranteed to be the global optimal solution. 7. Illustrative examples 7.1. A four-stage tensegrity tower 7.1.1. Initial and target configurations A tensegrity tower consisting of four stacked quadriprism modules named A, B, C, and D from the bottom to the top is shown in Fig. 2 . In the initial configuration, the bottom and top surfaces of each module are squares with side length a =2 \u221a 2 . For module A or C, the top surface is rotated through an angle of \u03b8 = 1/8 \u03c0 counterclockwise around the z -axis with respect to the bottom surface ( Fig. 3 (a)). The module B or D is rotated clockwise through the same angle ( Fig. 3 (b)). Modules B, C and D are rotated clockwise around the z -axis by 1/4 \u03c0 , 0, and 1/4 \u03c0 , respectively, relative to module A, as shown in Fig. 3 (c). Two adjacent modules are assem- b c m t b b n A m o s t h o m t g A r s h f a m d i 7 c f s a 1 m t t o \u2212 r x m 0 m 8 t b A i o o a t t m C b t c t i b b b t a T m i 1 led by replacing the cables in their shared surfaces with saddle ables. Diagonal cables are then added to the side surfaces of each odule. In addition, the side elements are arranged in opposite roation directions for two adjacent modules. There are 32 joints, 16 ars and 72 cables in total in the tensegrity tower. The joint numers are also shown in Fig. 2 , and each element is denoted by the umbers of its two end joints. Six DOFs, A1- x , A1- y , A1- z , A2- x , 2- z , A3- z , are constrained on the ground to prevent rigid-body otion of the tower. The height of each module is H = 8, and the overlap depth f adjacent modules is h . To be a tensegrity, a particular relationhip must be satisfied between \u03b8 and h / H ( Nishimura, 20 0 0 ). For he initial configuration shown in Fig. 2 , it can be determined that / H = 0.29509 corresponds to \u03b8 = 1/8 \u03c0 . The numbers of the modes f self-stress state and inextensional mechanism are s = 1 and = 3, respectively. The tensegrity tower is expected to deform from the above iniial configuration to a target configuration shown in Fig. 4 . The taret configuration is constructed by forcing the joints in surfaces -II, B-I, B-II, C-I, C-II, D-I and D-II to move in the negative di- ection of the x axis. The moving distance of joints in the same urface is d = kz 2 , where k is a prescribed constant, and z is the eight of the surface from the ground" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure7.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure7.5-1.png", "caption": "Figure 7.5 Primary and secondary clearance", "texts": [ " to cemented carbide, cermet has improved wear D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 7 Cutting tools and cutting fluids 105 7 The clearance angle should be kept at an absolute minimum, 8\u00b0 being quite adequate for most purposes. Grinding an excessive clearance angle should be avoided \u2013 it is a waste of expensive cutting-tool material, a waste of time and money in grinding it in the first place, and, finally and most important, it weakens the cutting edge. In some cases, however, a greater clearance angle may be required, e.g. where holes are being machined using a boring tool, Fig. 7.5(a). If this additional clearance is provided up to the cutting edge, Fig. 7.5(b), serious weakening will result, so it is customary in these instances to provide the usual clearance angle for a short distance behind the cutting edge, known as primary clearance, followed by a second angle known as a secondary clearance angle, Fig. 7.5(c). and high-speed rough machining cast iron by turning and milling operations. CBN is used in applications that require extreme wear resistance and toughness and is the only cutting material that can replace traditional grinding methods. CBN is referred to as a superhard cutting material. Polycrystalline diamond (PCD) is a composite of diamond particles sintered together with a metallic binder. Diamond is the hardest and therefore the most abrasion resistant of all materials. As a cutting-tool material, it has good wear resistance but lacks chemical stability at high temperatures", " A suitable lubricant should be used to enhance the life of the Figure 9.28 Boring tool A boring tool must be smaller than the bore it is producing, and this invariably results in a thin flexible tool. For this reason it is not usually possible to take deep cuts, and care must be taken to avoid vibration. In selecting a boring tool, choose the thickest one which will enter the hole, to ensure maximum rigidity. Ensure also that adequate secondary clearance is provided in relation to the size of bore being produced, as shown in Fig.\u00a07.5. Knurling is a process whereby indentations are formed on a smooth, usually round, surface to allow hands or fingers to get a better grip. Fig.\u00a09.29 shows diamond knurling on a tap wrench. Knurling can also be used purely for decorative purposes. The indentations are formed by pressing hardened knurling wheels (known as knurls) held in a knurling tool, Fig. 9.30, onto the workpiece and deforming the surface. The knurls may produce a straight, diagonal or crisscross (known as a diamond knurl) pattern on the workpiece, the form of which can be coarse, medium or fine depending on the knurls used" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001809_1.4033621-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001809_1.4033621-Figure1-1.png", "caption": "Fig. 1 Functional scheme of dry dual clutch system", "texts": [ " The main advantages of the electromechanical actuation solutions include easier integration into the whole transmission system and potentially better total efficiency due to low losses in electrical motors [8\u201311]. For the purpose of designing a clutch torque control system for a favorable DCT control performance during vehicle launch and gear shifting [12,13], it is valuable to have a good insight into the clutch system static and dynamic behavior, and to provide a mathematical model capable of accurately describing the system behavior. To this end, this paper proposes a dynamic model of a dry dual clutch system with lever-based electromechanical actuators illustrated in Fig. 1. The model is developed by using the bond graph methodology, which is a proven method for modeling and power flow analysis of multiphysical dynamic systems [14], including automotive systems and their components [15\u201318]. The main advantage of the bond graph method is the provision of a simple conversion of the multiphysical system (combining, e.g., electrical, mechanical, hydraulic, and elements from other physical domains) to a symbolic graph that mirrors the original system structure and is highly modular, which makes it easy to expand or modify", " Contributed by the Dynamic Systems Division of ASME for publication in the JOURNAL OF DYNAMIC SYSTEMS, MEASUREMENT, AND CONTROL. Manuscript received July 23, 2015; final manuscript received February 23, 2016; published online June 15, 2016. Assoc. Editor: Zongxuan Sun. Journal of Dynamic Systems, Measurement, and Control SEPTEMBER 2016, Vol. 138 / 091012-1 Copyright VC 2016 by ASME Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jdsmaa/935328/ on 03/04/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use Fig. 1, together with a definition of main system variables and parameters. 2.1 Actuator. The actuator is based on a lever which transfers the preloaded spring force Fs into the variable output force Flo, depending on the position xf of a fulcrum driven by an electric servodrive. Apart from the fulcrum and the engagement bearing, the lever motion is constrained by attaching it to a soft leaf spring, and by a stiff support placed at the lever back end. Consequently, during operation, the lever rotates by the angle al and translates by xlf in the radial direction (with respect to screw drive axis). Its motion is determined by a balance of forces and torques produced by a pair of preloaded helical springs (Fs), the engagement bearing (Flo), and the support (Fsup). Based on experimental findings [19], the lever is considered to be compliant (the parameter kl in Fig. 1), which accounts for the effects of lever, fulcrum, and engagement bearing elasticity. The actuator can generally operate in two distinct modes: (i) the lever is leaned to the support (mode 1, Fsup> 0) and (ii) see-saw mode (mode 2, Fsup\u00bc 0). The fulcrum itself represents an integral part of the system of rollers (actually a cart, see the right-hand side of Fig. 1), which is connected to the servomotor by means of a precise ball screwdrive with the speed ratio isd. The screw-drive is directly connected to the shaft with radius r12, around which the rollers 1 and 2 (with the radii r1 and r2) freely rotate. The roller 1 rolls over the base plate track, thus providing the radial support to the screwdrive. The roller 2 drives the fulcrum roller (with the radius r3) by means of direct contact. Finally, the roller 4 plays the role of a fulcrum roller shaft with the contact radius r34, and it rolls over the base plate track. Due to the relatively large fulcrum roller radius r3, there is an axial eccentricity lce(xf) of the lever-fulcrum contact point with respect to fulcrum roller center (Fig. 1). The fulcrum roller is loaded by the normal force FN, acting at the contact angle b(xf) al, where b is the lever profile angle that depends on the fulcrum position xf. The normal force radial component FNz is transferred to the base plate track as the load of roller 4. At the same time, the normal load axial component FNx is transferred to the roller 2, and through the rollers 1/2 shaft to the screw-drive as its load. In addition to FNx, the total screw-drive axial force Fx includes the resistive forces caused by the rolling resistance of rollers 3 and 4 (denoted as Mfr3 and Mfr4), and sliding friction at the radii r34 (Mf34) and r12 (Mf12). The friction torques Mf34 and Mf12 depend on the fulcrum normal load radial component FNz and the screw-drive load Fx, respectively. The rolling resistance and sliding friction of roller 1 are negligible due to a small normal force for that roller. 2.2 Dual Clutch Assembly. The dual clutch assembly (the left-hand side of Fig. 1) comprises the flywheel (a center plate), two pressure plates with corresponding leaf-type return springs, and a moveable cover attached to pressure plate 1, two diaphragm springs with associated clutch wear compensation mechanisms, and two friction plates. The friction plates are connected to the corresponding transmission input shafts by spline coupling, and they can move axially. All other components are connected to the flywheel, which is attached to the engine crankshaft. The connection between the pressure plates and the flywheel is established through the leaf return springs", " The friction plates are of similar design as the ones used in conventional manual transmission clutches (see, e.g., Ref. [20]). They consist of a hub, an elastic core plate (cushion plate), and friction linings. The elastic core plate provides smooth engagement and uniform contact between the friction linings and the mating surface, i.e., the pressure plates and the flywheel. Figure 2 describes the way in which the clutch assembly is modeled (for clutch 2) and how it interacts with the actuator. Here, it should be noted that each clutch of the dual clutch assembly (Fig. 1) can be considered separately, because there is no mechanical interaction between them. In order to derive a simple, control-oriented model that still has an accurate input\u2013output behavior, several assumptions are taken into account: (i) the diaphragm spring is assumed to be leaned against the housing (while in reality, it is supported by the pressure plate), (ii) friction in different dry contact points of the diaphragm spring levers is lumped into a single input-side friction element (Ff.ds), and (iii) the compliance of different components is lumped into a single, friction plate-side compliance (1/kfp)", " Several custom-made test rigs have been designed and used for clutch subsystems\u2019 experimental characterization and model validation. Figure 3 shows the schematics of these rigs, while the corresponding photographs and video clips can be found in Refs. [21] and [22], together with more detailed information on the test rig design. Figure 3(a) illustrates the structure of clutch actuator test rig, which is built around the original clutch bell housing that accommodates the electromechanical actuators (see the top photograph in Fig. 1). The actuator is driven by a general-purpose permanentmagnet synchronous servomotor, fitted with a precise incremental encoder. Using a linear helical spring as the actuator, load facilitates detailed characterization of the actuator dynamics, while avoiding the influence of dual clutch assembly nonlinearities and their interaction with the actuator dynamics. Three helical springs with different stiffness have been employed for testing the actuator under different loading conditions. Similarly, the rig allows for emulating an initial actuator preload or clearance by means of a manually operated adjusting mechanism", " Also, both static modeling approaches have difficulties with including the actuator friction effects due to the additional system coupling paths introduced by friction. The disadvantages of the static model formulations have been overcome by proposing a dynamic actuator model, which introduces the lever inertia dynamics (bond 26) in order to decouple the input and output subsystems (no algebraic loops) and describe the friction effects in a physical manner. The lever dynamics submodel realization is based on the equilibrium of forces and torques acting on the lever in the two system degrees-of-freedom (xlf and al, see Fig. 1). Here, the screw-drive velocity vx is transformed to the lever subsystem input velocity vlf through the gain of modulated transformer (MTF) placed between bonds 18 and 19. The MTF gain i 1 l \u00bc tan \u00bdb\u00f0xf \u00de al \u00fe atan \u00f0xl=\u00f0lb0 lcf \u00f0xf \u00de\u00de\u00de (1) is predominantly related to the lever-fulcrum contact angle b(xf) al (see Fig. 1), and it also accounts for the lever bending angle calculated from the lever radial deflection xl. The lever subsystem bond graph model in Fig. 4 directly reflects the structure of the mechanical system in Fig. 1. The lever radial/lift velocity vlf, together with the corresponding displacements due to lever rotation with speed xl, drives the three distinct actuator compliance elements (energy spring, support, and output compliance bonds 21, 31, and 34), in order to calculate the radial component FNz of the total lever load FN. The lever dynamics model is extended with the sliding friction torque element 27, and an artificially introduced linear damping element 28 tuned to suppress oscillations of the over-dimensioned lever inertia to a numerically acceptable level", " 4 gives the following set of dynamic equations: Jm _xm \u00bc Mm i 1 sd i 1 l \u00f0Flo \u00fe Fs Fsup\u00de \u00bdMf m \u00feMfsd \u00fe i 1 sd \u00f0r 1 2 Mf 12 \u00fe r 1 3 Mfr3 \u00feMf 34\u00f0r 1 3 \u00fe r 1 4 \u00de \u00fe r 1 4 Mfr4\u00de |fflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl{zfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflfflffl} Mfm:e (2) Jl _xl \u00bc Fs\u00f0ls \u00fe lcf \u00f0xf \u00de\u00de Flo\u00f0lb lcf\u00f0xf \u00de\u00de Fsup\u00f0lsup \u00fe lcf \u00f0xf \u00de\u00de blxl Mf l (3) mpp _vpp \u00bc ids\u00f0Flo Fsup:b Fds Ff :ds\u00de Frs Fn (4) _Fs \u00bc ks\u00bdi 1 l i 1 sd xm \u00f0ls \u00fe lcf \u00f0xf \u00de\u00dexl (5) _Fsup \u00bc ksup\u00bd\u00f0lsup \u00fe lcf \u00f0xf \u00de\u00dexl i 1 l i 1 sd xm (6) _Flo \u00bc kl\u00bdi 1 l i 1 sd xm \u00fe \u00f0lb lcf \u00f0xf \u00de\u00dexl idsvpp (7) _Fds \u00bc kdsidsvpp (8) _Frs \u00bc krsvpp (9) _Fn \u00bc kfpvpp (10) _Fsup:b \u00bc ksup:bidsvpp (11) where the lever transformer ratio 1/il is defined by Eq. (1), and where most of the stress\u2013strain curves represented by the stiffness coefficients k(.) in Eqs. (5)\u2013(11) are nonlinear (see Sec. 5). The basic dimensions and kinematic parameters of the actuator and clutch assembly mechanisms (Fig. 1) were measured using common workshop and laboratory tools such as digital calipers and a drawing compass. Other parameters were estimated based on insights into mechanism design or reconstructed from the test rig recordings, as described in Sec. 5. 5.1 Actuator Model 5.1.1 Actuator Drive. The actuator motor torque Mm, as the overall model input (Fig. 4), is generated by the same cascade controller of motor position am as the one used in the experimental system. A cogging torque component is added to the motor torque in order to include realistic high-frequency torque perturbations, and more importantly to provide realistic ditherlike conditions [24] for accurate friction modeling" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000984_asjc.1196-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000984_asjc.1196-Figure4-1.png", "caption": "Fig. 4. Scheme of the inverted pendulum.", "texts": [ " The structure for the optimization matrix-valued functions \u03a0(\u22c5) and L(\u22c5) is the same as considered for P1(\u22c5) in (23). To check, for instance, the satisfaction of the robust IO-FTS requirement, Fig. 3 reports the weighted output zT Qz for 100 random realizations of the uncertain matrices, when the exogenous input w(t) = 0.115, with t \u2208 [0 , 15], is considered. As expected, zT Qz is always less than 1. Furthermore, the \u221e bound implies \u2016z\u20162 ,\u03a3 < 1; in particular for the considered 100 random simulations it is \u2016z\u20162 ,\u03a3 < 0.66. As a case study, in this section the inverted pendulum shown in Fig. 4 is considered. Given the parameters reported in Table I, letting x = (s s\u0307 \ud835\udf11 ?\u0307?)T Table I. Parameters of the inverted pendulum. s cart position s\u0307 cart velocity \ud835\udf11 angular position of the pendulum ?\u0307? angular velocity related to \ud835\udf11 u control force applied to the cart w disturbance force applied to the pendulum M mass of the cart 0.5 Kg m mass of the pendulum 0.2 Kg b coefficient of friction of the cart 0.1 N\u2215(m \u22c5 sec) L length to the pendulum centre of mass 0.3 m I mass moment of inertia of 0.006 Kg \u22c5 m2 the pendulum \u00a9 2015 Chinese Automatic Control Society and John Wiley & Sons Australia, Ltd 838 Asian Journal of Control, Vol" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002274_j.mechmachtheory.2019.103748-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002274_j.mechmachtheory.2019.103748-Figure10-1.png", "caption": "Fig. 10. 3-UPU representative limb, link labels are shown between parentheses.", "texts": [ " Additional details needed for motion and action analyses include the position of the couplings and the orientation of the axes, all unambiguously written with respect to a reference frame. The reference frame is located with the origin coincident to the centre of equilateral triangle formed by the universal joint centres connected to the base, with the z -axis orthogonal to the base plane and the x -axis pointing to a universal joint centre (see Figs. 11 and 13 ). This specific universal joint belongs to what is called the first limb. Since all limbs are identical, although placed at different locations, only this first limb, which is shown schematically in Fig. 10 , is described herein. In Fig. 10 , the coupling labelled A, of the universal kind, connects the base, labelled 0, to link 1. Link 1 is connected to link 2 by a prismatic joint modelled by the couplings B and C (for more details see Section 5.2 ). Finally, link 2 is connected to link 3, the platform, by another universal joint labelled D. Point A is located at the universal joint centre with the same label. Link 2 stands on link 1 at point B and the distance between A and B is constant. A similar situation is found for points C and D. The location of the base-connected universal joint centre (e.g., point A in Fig. 10 ) of each limb with respect to the reference frame is given by the position vector e1 = [ x cos \u03b1 y sin \u03b1 z 0 ] r b (35) where r b is the distance from the base centre (the origin) to any universal joint centre connected to the base, and \u03b1 = 0 , 2 \u03c0/ 3 , \u22122 \u03c0/ 3 rad for first, second and third limbs, respectively. The orientation of the first axis of every universal joint, see Fig. 10 , is given by the unity vector \u02c6 e1 of e1 , since the platform rotation is prevented. The vector pointing from the universal joint centre of the base to that of the platform for each limb is given by e2 = [ x M x + ( r t \u2212 r b ) cos \u03b1 y M y + ( r t \u2212 r b ) sin \u03b1 z M z ] (36) where r t is the distance from the platform centre to any universal joint centre connected to the platform; M x , M y , and M z are the coordinates of the platform centre (point M) with respect to the reference frame; and r b and \u03b1 are defined for Eq", " The direction (axis) of the prismatic joint of each limb is given by the unity vector \u02c6 e2 of e2 . The orientation of the second axis of every universal joint is given by the unity vector \u02c6 e3 , orthogonal to \u02c6 e1 and \u02c6 e2 . Couplings B and C transmit forces parallel to \u02c6 e3 and \u02c6 e4 ; \u02c6 e4 is orthogonal to \u02c6 e2 and \u02c6 e3 . The universal joints transmit a torque around the unity vector \u02c6 e5 , orthogonal to \u02c6 e1 and \u02c6 e3 . The five unity vectors \u02c6 ei , ( i = 1 , 2 , . . . 5 ) are dependent on the location of the respective limb given by \u03b1. For the first limb, \u03b1 = 0 , the \u02c6 ei are shown in Fig. 10 . The location of the universal joint centres on the base (i.e. points A, E, and I) are given by e1 , from Eq. (35) , with a suitable angle \u03b1. Also, the location of the universal joint centres on the platform (i.e. points D, H, and L) are given by e1 + e2 , from Eqs. (35) and (36) . Neglecting the radial length of the prismatic joints, the two contact points between links 1 and 2 are given by B = e1 + l B \u02c6 e2 C = e1 + e2 \u2212 l C \u02c6 e2 (37) where l B is the constant distance between points A and B, and l C is that of points C and D" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001447_asjc.1165-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001447_asjc.1165-Figure1-1.png", "caption": "Fig. 1. Partition of the state space.", "texts": [ "2 Piecewise Lyapunov functions \u039bi \u2236= \u23a7\u23aa\u23a8\u23aa\u23a9 [ sin \ud835\udf03i \u2212 cos \ud835\udf03i ] [ uc yc ] \u2265 0[ sin \ud835\udf03i\u22121 \u2212 cos \ud835\udf03i\u22121 ] [ uc yc ] \u2264 0 , \u039b\ud835\udf00 \u2236= \u23a7\u23aa\u23a8\u23aa\u23a9 [ sin \ud835\udf030 \u2212 cos \ud835\udf030 ] [ uc yc ] \u2265 0[ sin \ud835\udf03\ud835\udf00 \u2212 cos \ud835\udf03\ud835\udf00 ] [ uc yc ] \u2264 0 , \u039b0 \u2236= \u23a7\u23aa\u23a8\u23aa\u23a9 [ sin \ud835\udf03\ud835\udf00 \u2212 cos \ud835\udf03\ud835\udf00 ] [ uc yc ] \u2265 0[ sin \ud835\udf03N \u2212 cos \ud835\udf03N ] [ uc yc ] \u2265 0 (14) Note that the obtained LMI conditions in Theorem 1 have some limitations because of the conservativeness of using quadratic terms in Lyapunov functions. In this section, we use piecewise Lyapunov functions such that less conservative stability conditions can be obtained. Based on (5) we have [ uc yc ] = [ \u2212Cp 0 0 Cc ] \u23df\u23de\u23de\u23de\u23de\u23df\u23de\u23de\u23de\u23de\u23df T [ xp xc ] = Tx. The flow set and jump set defined in (8) are shown in Fig. 1, the state space is partitioned into N +2 angular sectors. Let N \u2265 1, define \ud835\udf03i = \ud835\udf0bi 2N , i = {0, 1, \u00b7 \u00b7 \u00b7 ,N} such that 0 = \ud835\udf030 < \ud835\udf031 < \u00b7 \u00b7 \u00b7 < \ud835\udf03N = \ud835\udf0b\u22152, \ud835\udf03\ud835\udf00 = arctan(\u2212\ud835\udf00\u22152). The sector \u039bi, i = {\ud835\udf00, 0, 1, \u00b7 \u00b7 \u00b7 ,N} can be represented as the form of (14). Denote \u0398i = [ sin \ud835\udf03i \u2212 cos \ud835\udf03i ]T . Based on [ uc yc ] = Tx and define the following matrices Si \u2236= \u2212TT (\u0398i\u0398T i\u22121 + \u0398i\u22121\u0398T i ) T , i = {1, \u00b7 \u00b7 \u00b7 ,N} S\ud835\udf00 \u2236= \u2212TT (\u03980\u0398T \ud835\udf00 + \u0398\ud835\udf00\u0398T 0 ) T S0 \u2236= TT (\u0398\ud835\udf00\u0398T N + \u0398N\u0398T \ud835\udf00 ) T Then, the sectors in (14) can be represented as \u039bi \u2236= { x \u2208 R n|xT Six \u2265 0 } , i = {\ud835\udf00, 0, 1, \u00b7 \u00b7 \u00b7 ,N} \u00a9 2015 Chinese Automatic Control Society and Wiley Publishing Asia Pty Ltd Denote \u0398i\u22a5 as the orthogonal matrices of \u0398i, where \u0398T i\u22a5\u0398i = 0,\u0398i\u22a5 = [ cos \ud835\udf03i sin \ud835\udf03i ]T ", " For given h > 0, \ud835\udf07 > 0, \ud835\udefc > 0, if there exist symmetric positive definite matrices Pi \u2208 Rn\u00d7n,Xi \u2208 R(2n+1)\u00d7(2n+1), i = {\ud835\udf00, 1, \u00b7 \u00b7 \u00b7 ,N},P0,Q,R \u2208 Rn\u00d7n, matrices Yi1 \u2208 Rn\u00d7n, Yi2 \u2208 Rn\u00d7n,Yi3 \u2208 R, i = {\ud835\udf00, 1, \u00b7 \u00b7 \u00b7 ,N}, constants \ud835\udf0fi \u2265 0, i = {\ud835\udf00, 0, 1, \u00b7 \u00b7 \u00b7 ,N}, \ud835\udefe > 0, 0 < \ud835\udeff < 1 such that the following conditions are satisfied \u23a1\u23a2\u23a2\u23a2\u23a3 \u03a8i 11 + \ud835\udefcPi + \ud835\udf0fiSi \u03a8i 12 \u03a8i 13 CT \u2217 \u03a8i 22 hAT d RB \u2212 Y T i3 0 \u2217 \u2217 hBT RB \u2212 \ud835\udefe 0 \u2217 \u2217 \u2217 \u2212\ud835\udefe \u23a4\u23a5\u23a5\u23a5\u23a6 + [ hXi 0 0 0 ] \u2264 0, i = {\ud835\udf00, 1, \u00b7 \u00b7 \u00b7 ,N} (15a)[ Xi [ Y T i1 Y T i2 Y T i3 ]T \u2217 e\u2212\ud835\udefchR ] \u2265 0, i = {\ud835\udf00, 1, \u00b7 \u00b7 \u00b7 ,N} (15b) AT r P0Ar \u2212 P0 \u2212 \ud835\udf0f0S0 < \u2212\ud835\udeffP0 (15c) \u0398T i\u22a5T\u2212T (Pi \u2212 Pi+1)T\u22121\u0398i\u22a5 = 0, i = {1, \u00b7 \u00b7 \u00b7 ,N \u2212 1} (15d) \u0398T 0\u22a5T\u2212T (P\ud835\udf00 \u2212 P1)T\u22121\u03980\u22a5 = 0 (15e) \u0398T \ud835\udf00\u22a5 T\u2212T (P0 \u2212 P\ud835\udf00)T\u22121\u0398\ud835\udf00\u22a5 = 0 (15f) \u0398T N\u22a5 T\u2212T (PN \u2212 P0)T\u22121\u0398N\u22a5 = 0 (15g) where \u03a8i 11,\u03a8 i 12,\u03a8 i 13,\u03a8 i 22 are defined as follows \u03a8i 11 = PiA + AT Pi + Q + Yi1 + Y T i2 + hAT RA \u03a8i 12 = PiAd \u2212 Yi1 + Y T i2 + hAT RAd \u03a8i 13 = PiB + Y T i3 + hAT RB \u03a8i 22 = \u2212(1 \u2212 \ud835\udf07)e\u2212\ud835\udefchQ \u2212 Yi2 \u2212 Y T i2 + hAT d RAd (16) Then, there exists a small enough \ud835\udf0c\u2217 > 0 such that for any \ud835\udf0c \u2208 (0, \ud835\udf0c\u2217], the origin of x dynamics is exponentially stable, and the system is finite gain 2 stable from w to y with the 2 gain upper-bounded by ?\u0304? = \ud835\udefe + \ud835\udf16 for any \ud835\udf16 > 0. Proof. Choose piecewise Lyapunov functions as follows V (xt) = xT t Pixt + \u222b t t\u2212d(t) xT (s)e\ud835\udefc(s\u2212t)Qx(s)ds + \u222b 0 \u2212h \u222b t t+\ud835\udf03 x\u0307T (s)e\ud835\udefc(s\u2212t)Rx\u0307(s)dsd\ud835\udf03, for xt \u2208 \u039bi, i = {\ud835\udf00, 0, 1, \u00b7 \u00b7 \u00b7 ,N} (17) Without loss of generality, consider the patching surface between \u039b1 and \u039b2 (Fig. 1), for (uc, yc) on this surface, we have the representation (uc, yc) = l(cos \ud835\udf031, sin \ud835\udf031), l \u2208 R. Then, from [ uc yc ]T = Tx, it follows that x = T\u22121 [ uc yc ]T = lT\u22121 [ cos \ud835\udf031 sin \ud835\udf031 ]T = lT\u22121\u03981\u22a5 on this surface (where T\u22121 is the left inverse of T), based on (15c), it follows that xT P1x = xT P2x, x \u2208 \u039b1 \u2229 \u039b2. Hence, the form of Lyapunov function (17) and (15c)\u2013(15f) imply that the Lyapunov function (17) is continuous on the patching surface. Then, take similar steps as in the proof of Theorem 1, it can be derived that the origin of x dynamics is exponentially stable, and the system is finite gain 2 stable from w to y with the 2 gain upper-bounded by " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001268_s0263574714002653-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001268_s0263574714002653-Figure1-1.png", "caption": "Fig. 1. (Colour online) The translational parallel manipulator RAF.", "texts": [ " The paper is organized as follows: in Sections 2 and 3, the inverse kinematic and the static models are developed. In Section 4, the model that determines the orientation error of the translational parallel manipulator RAF, using a deterministic method, is presented in an analytical form. The distribution of the orientation error of the manipulator caused by the joints in the two passive kinematic legs (PKLs) is presented. Some concluding remarks are presented in Section 6. The translational parallel manipulator RAF proposed by Romdhane et al24 (as shown in Fig. 1) is composed of a platform connected to the base by three active legs of type SPS (S and P stand for spherical and active prismatic pairs, respectively) and two PKLs, which are used to eliminate all possible rotations of the platform with respect to the base. Each PKL is composed of an arm (2) connected to the base by a revolute pair with axis i, i = 1 for the first PKL and i = 2 for the second PKL (as shown in Figs. 2 and 3). Two forearms (3) and (3\u2032) are connected to the arm (2) by two spherical joints centered at C and C\u2032" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000827_1077546315590908-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000827_1077546315590908-Figure4-1.png", "caption": "Figure 4. The rotor-bearing system. (a) Photograph of the test rig and (b) Schematic diagram of the test rig.", "texts": [ " ~Z R i \u00bc TnTn 1 T2T1\u00f0 \u00de ~Z L i \u00f07\u00de At two ends of the shaft, free boundary conditions are: ML 1 \u00bc 0, QL 1 \u00bc 0, MR n \u00bc 0, QR n \u00bc 0 \u00f08\u00de The transfer relation is: y M Q 1 8>>>>>>< >>>>>>: 9>>>>>>= >>>>>>; R n \u00bc T11 T12 T13 T14 T15 T21 T22 T23 T24 T25 T31 T32 T33 T34 T35 T41 T42 T43 T44 T45 0 0 0 0 1 2 6666664 3 7777775 n y M Q 1 2 6666664 3 7777775 L 1 \u00f09\u00de y L 1 \u00bc T31 T32 T41 T42 1 T35 T45 \u00f010\u00de The state vector ~Z L 1 can be obtained from equation (10). Then the coupling state vector is obtained from the transfer matrix equation. An experimental system is established for verifying the strain gauge method, as shown in Figure 4(a). The shaft system consists of 3 rotors, 2 couplings and 6 bearings. Each rotor is supported by two journal bearings at the two ends. There are 16 discs with different diameters on the shaft. The total shaft weight is 5420N. Figure 4(b) gives the corresponding schematic diagram of the test rig and the strain gauge locations. Jack-up load tests are also conducted to provide an independent verification of the strain gauge alignment model. Besides, the laser alignment method is used to measure the sag/gap values of couplings for comparison. The relative error Err is defined as: Err \u00bc A Bj j A\u00feB 2 100% \u00f011\u00de where A is the identification value (bearing load, sag or gap) of the first identification method (jack-up method, at UNIVERSITE DE MONTREAL on September 5, 2015jvc" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002627_0954406220945728-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002627_0954406220945728-Figure6-1.png", "caption": "Figure 6. Analysis of the joint interface between B- and Z-axis.", "texts": [ " So new kinematic parameters can be defined by the displacement expression of the free end, which can be expressed only by the motion amount of A-axis and the new kinematic parameters zB \u00bc Gh EBAB LB \u00bc g1 \u00f014\u00de yB \u00bc GhLh LB \u00fe zB\u00f0 \u00de 2 sin 2EBIB \u00bc g2 sin \u00f015\u00de XB \u00bc GhLh LB \u00fe zB\u00f0 \u00de sin EBIB \u00bc g3 sin \u00f016\u00de The new kinematic parameters g1, g2, g3\u00bd are introduced in the analysis of the joint interface between A- and B-axis, and they will be calculated in subsequent section of parameter identification. Analysis of the joint interface between B- and Z-axis The deformation of the joint interface between B- and Z-axis is affected both by motion of A- and B-axes, as shown in Figure 6(a), where the blue line represents elastic beam deformation model and its length is LZ, the elastic modulus is EZ, the cross-sectional area is AZ, and the moment of inertia are IXZ and IYZ. In addition, the motions of the A- and B-axes are a, . The analysis of the elastic beam is shown in Figure 6(b), and the deformation at the free end of the elastic beam is zZ \u00bc Gh EZAZ LZ \u00bc g4 \u00f017\u00de xZ \u00bc GhLh LZ \u00fe zZ\u00f0 \u00de 2 cos sin 2EZI Y Z \u00bc g5 cos sin \u00f018\u00de yZ \u00bc GhLh LZ \u00fe zZ\u00f0 \u00de 2 sin 2EZI X Z \u00bc g5 sin \u00f019\u00de XZ \u00bc GhLh LZ \u00fe zZ\u00f0 \u00de sin EZI X Z \u00bc g6 sin \u00f020\u00de YZ \u00bc GhLh LZ \u00fe zZ\u00f0 \u00de cos sin EZI Y Z \u00bc g7 cos sin \u00f021\u00de Similarly, four new kinematic parameters are introduced in the analysis of the joint interface between Band Z-axis: g4, g5, g6, g7\u00bd , and corresponding gravity deformation matrix is Z BTg \u00bc c YZ 0 s YZ xZ 0 1 0 0 s YZ 0 c YZ zZ 0 0 0 1 2 6664 3 7775 1 0 0 0 0 c XZ s XZ yZ 0 s XZ c XZ 0 0 0 0 1 2 6664 3 7775 \u00bc c YZ s XZs Y Z c XZs Y Z xZ 0 c XZ s XZ yZ s YZ s XZc Y Z c XZc Y Z zZ 0 0 0 1 2 6664 3 7775 \u00f022\u00de The analysis of the joint interface between Z- and Y-axis is shown in Figure 7(a), and the analysis of the joint interface between Y- and X-axis is shown in Figure 7(b), where blue lines represent elastic beam deformation models and are fixed at one end and free at the other end" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001221_esda2014-20001-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001221_esda2014-20001-Figure4-1.png", "caption": "FIGURE 4. MECHANICAL CONTROL ELEMENTS OF THE IN-SITU LASER CLADDING DEVICE", "texts": [ " The device also comprises two upper rods, positioned in the upper part of the frame part, and two lower rods in the lower part of the frame part, by means of which both frame parts are rigidly connected to each other (see Fig. 3). The device further comprises two carriages which are installed on the upper rods and lower rods so that both carriages can be slid along these rods. A laser nozzle is installed operatively between both carriages. The device includes two servo motors, the first of which is installed in the first carriage and operatively connected to the laser nozzle to control its pivoting angle (see Fig. 4). A second servo motor is installed on the second carriage and is operatively connected to one of the two lower rods by means of a gearing transmission to control the laser nozzle\u2019s longitudinal position. The device\u2019s gearing transmission comprises a spur gear connected to the second servo motor and a toothed rack, which is formed directly by the lower rod. The first servo motor is connected to the laser nozzle through a bush (see Fig. 5). To ensure the device\u2019s positioning and controlled rotation around the crankshaft journal, it comprises the two aforementioned guide-ways and two opposite-guide-ways" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001347_j.ymssp.2014.06.016-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001347_j.ymssp.2014.06.016-Figure2-1.png", "caption": "Fig. 2. Mechatronic design process diagram.", "texts": [ " While the control signal and dynamics of the nonlinear mobile platform is performed in real-time, the embedded system is striving significantly for the CPU computing power. In this stage, time extended mobile platform operability (electrical power), together with error movement level and controller performance, should be equilibrated. Consequently the mechatronic attitude incorporated into the presented development methodology is complemented with optimisation design loops, facilitating the balanced design. It is portrayed a complementary perspective of the design flow (Fig. 2). Segregated stages of the adopted mechatronic process merge optimisation feedback loops, enabling during iterative cycles for the optimal system design. The detailed overview of the applied optimisation loops in following study was summarized in Table 1. There are several not commonly used features that distinguish the presented mobile platform design among the other constructions of the same type. Propulsion is given by electrical motor coupled on the same shaft with differential fitted into mobile platform chassis", " determine the length of integration step (system command frequency) ensuring that the operation of the control system is executed in real-time at which the errors occurrence fulfils satisfactory level. Appropriate selection of integration step followed by a feedback loop: virtual prototyping technique\u2014the HILS technique. The controller obeys real time performance if all the foreground processes (algorithm related) have been successfully executed before their stated critical time constraints. The authors, by the choice of step length \u0394t verified equalization of computation time (system time) with actual time. The considered optimisation loop was depicted in Fig. 2. Provisory input for step length analysis can flow in when the deviation levels measured between desired torque trajectories and trajectories obtained in the process of virtual prototyping technique are acceptable. During the virtual prototyping technique (first stage of this study), the authors explored the \u0394t step sets, but demanding computation process of on-line algorithm and real-time conditions (equalization between simulation and real time clock readings) of the HILS tests verified authors assumptions and final \u0394t had to be set for 0", " Due to the cumbersome system description, the calculation process of nonlinear equations, implemented the LabVIEW algorithm, absorbs significant computer resources (during the calculations the processor is loaded with approximately 100%). Therefore, in order to achieve feasible solution within equipment available performance (the type of available controller was known before the main phase of study), there was extremely important to perform optimisation process (i.e. surveillance system energy optimisation, Fig. 2). It should be noted that the system optimisation was obtained during iteration processes. The optimisation process is: associated with code architecture of deployed algorithm, and the place of its execution to secure request for controller performance (it was decided to separate the part of algorithm executed in the Real-Time processor and the part executed in the FPGA unit), energetic (the most noticeable) from a point of view of the control system (i.e. the choice of integration step). The applied virtual prototyping technique lowered considerably the risks associated with the design, by improving understanding of the system requirements" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002452_rspa.2020.0062-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002452_rspa.2020.0062-Figure6-1.png", "caption": "Figure 6. Lagrangian variables for the deformed beam. (a) The deformed state \u03d5i , i = 1 . . . n and equilibrium on the i-th segment; (b) perturbation\u03d5i + \u03b4\u03d5i from the deformed state.", "texts": [ " In the following, we will limit ourselves to the case in which the pitch lines are symmetric with respect to both the tangent and the normal axes at their initial pitch point P0. In this condition, the second couple of conjugate profiles can be obtained by mirroring the first couple with respect to the tangent axis of symmetry (figure 3). In this case, all the segments composing the beam, except at most the end ones, are identical: the assembly is done by simply rotating half the segments by 180\u25e6 about their longitudinal axis, so as to bring the conjugate profiles in contact, as indicated in figure 4a. (b) The equilibrium equations With reference to figure 6a, the absolute rotations \u03d5i of the segments i, i = 1 . . . n, positive if clockwise, are chosen as the Lagrangian variables to describe the deformation of the beam in figure 2. Since the segments are assumed rigid, the elastic strain energy U is due to the deformation of the cable only. Let K represent the elastic constant of the cable as a spring2 and denote with \u039b0 the cable elongation in the reference state of a straight beam, so that the initial pre-tension is T0 = K \u039b0. If \u039b is the successive variation of the cable length due to the inflection of the beam, the elastic strain energy in the cable is U = \u222b\u039b+\u039b0 \u039b0 K\u039b\u2032 d\u039b\u2032 = K\u039b0\u039b + 1 2 K\u039b2", "A476:20200062 ........................................................... Clearly, K(\u039b0 + \u039b) is the tensile force T in the cable, uniform along its length from the hypothesis of negligible friction. Whereas \u039b0 is a datum from the initial prestress, \u039b =\u2211n\u22121 i=1 \u039bi is a nonlinear function of all \u03d5i and depends upon the selected shape for the pitch lines at the contact sections. Observe that the Langrange multiplier \u03bc plays the role of the upward vertical constraint reaction at the right-hand side roller of figure 6a, whereas equation (2.7)1 comes from the principle of virtual displacement for a variation \u03d5i + \u03b4\u03d5i at the i-th segment, as shown in figure 6b. In fact, one has that the corresponding variation v\u0304j + \u03b4v\u0304j is \u03b4v\u0304j = \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 0, for j < i, ( i \u2212 bi) cos(\u03d5i)\u03b4\u03d5i , for j = i, i cos(\u03d5i)\u03b4\u03d5i , for j > i. (2.8) In particular, (2.7)1 corresponds to the rotational equilibrium of the i-th segment about the ith node, indicated at the top of figure 6a. In fact, the bending moments at the (i \u2212 1)-th and i-th nodes are equal to Mi\u22121 = ai\u22121T and Mi = aiT, respectively, whereas the vertical shear force at node (i \u2212 1) is equal to Vi\u22121 = \u03bc \u2212\u2211n j=i Fj. Consequently, one has Mi \u2212 Mi\u22121 + Vi\u22121 i cos(\u03d5i) + Fibi cos(\u03d5i) = 0, (2.9) which clearly coincides with (2.7)1. It is important to observe from (2.7) that the moment Mi at the i-th spring hinge is a function not only of its relative rotation \u03d5i = \u03d5i \u2212 \u03d5i+1 but also of all the relative rotations \u03d5i for i = 1 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003429_j.tws.2021.107585-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003429_j.tws.2021.107585-Figure5-1.png", "caption": "Fig. 5. Edge loads for the pressure vessel with ellipsoidal ends.", "texts": [ " The boundary conditions applied on the displacement functions vary, depending on whether the closed apex or open apex shell is considered (Fig. 4). In the case of closed apex i.e. \ud835\udf111 = 0 following boundary conditions are assumed: ( ) ( ) \ud835\udc51\ud835\udc63 \ud835\udf111 = 0, \ud835\udc51\u210e \ud835\udf112 = 0, \ud835\udf171(\ud835\udf111) = 0. (40) The boundary conditions provided in Eqs. (40)\u2013(42) consider separated shell structures. To achieve structural compatibility in the junctions of deformed pressure vessel additional boundary conditions have to be applied. The compatibility equations i.e. boundary conditions for the junction of the ellipsoidal and cylindrical shell (Fig. 5) have the form: \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc522 ) = \ud835\udc51\ud835\udc63 ( \ud835\udf09\ud835\udc502 ) , \ud835\udf171 ( \ud835\udf11\ud835\udc522 ) = \ud835\udf171 ( \ud835\udf09\ud835\udc502 ) . (43) For the pressure vessel with torispherical dished ends (Fig. 6) three shells constituting two junctions have to be considered, therefore: \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc602 ) = \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc611 ) , \ud835\udf171 ( \ud835\udf11\ud835\udc602 ) = \ud835\udf171 ( \ud835\udf11\ud835\udc611 ) , \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc612 ) = \ud835\udc51\ud835\udc63 ( \ud835\udf09\ud835\udc502 ) , \ud835\udf171 ( \ud835\udf11\ud835\udc612 ) = \ud835\udf171 ( \ud835\udf09\ud835\udc502 ) . (44) After applying the boundary conditions described in Eqs. (40)\u2013(42) to the displacement functions it is possible to formulate the total potential energy as a function of the remaining \ud835\udc4e\ud835\udc56 and \ud835\udc4f\ud835\udc56 parameters", " (49) The above approach for obtaining displacement functions assumes that the edge loads contained in \ud835\udc53 (Eq. (48)) are known. The aim of this study is to take into consideration complex shell structures, i.e.: the structures that consist of more than one shell described by continuous principal radii of curvature. To analyse the interaction between connected shells it is necessary to formulate the compatibility equations based on Eqs. (43), (44). The edge loads for the pressure vessel with ellipsoidal dished ends are presented in Fig. 5. The compatibility equations are following: \ud835\udeff\ud835\udc501( \ud835\udc4b1)\ud835\udc4b\ud835\udc501 + \ud835\udeff\ud835\udc501( \ud835\udc4c 1)\ud835\udc4c\ud835\udc501 + \ud835\udeff\ud835\udc501( \ud835\udc4d1)\ud835\udc4d\ud835\udc501 + \ud835\udeff\ud835\udc501( \ud835\udc5d)\ud835\udc5d = \ud835\udeff\ud835\udc522( \ud835\udc4b2)\ud835\udc4b\ud835\udc522 + \ud835\udeff\ud835\udc522( \ud835\udc4c 2)\ud835\udc4c\ud835\udc522 + \ud835\udeff\ud835\udc522( \ud835\udc4d2)\ud835\udc4d\ud835\udc522 + \ud835\udeff\ud835\udc522( \ud835\udc5d)\ud835\udc5d, \ud835\udf17\ud835\udc501( \ud835\udc4b1)\ud835\udc4b\ud835\udc501 + \ud835\udf17\ud835\udc501( \ud835\udc4c 1)\ud835\udc4c\ud835\udc501 + \ud835\udf17\ud835\udc501( \ud835\udc4d1)\ud835\udc4d\ud835\udc501 + \ud835\udf17\ud835\udc501( \ud835\udc5d)\ud835\udc5d = \ud835\udf17\ud835\udc522( \ud835\udc4b2)\ud835\udc4b\ud835\udc522 + \ud835\udf17\ud835\udc522( \ud835\udc4c 2)\ud835\udc4c\ud835\udc522 + \ud835\udf17\ud835\udc522( \ud835\udc4d2)\ud835\udc4d\ud835\udc522 + \ud835\udf17\ud835\udc522( \ud835\udc5d)\ud835\udc5d. (50) The above formulae defines that in the deformed structure, displacements as well as rotations on the shells edges must be equal. Coefficients \ud835\udeff and \ud835\udf17 refer to the vertical displacement (Eq. (17)) and rotations (Eq. (18)) at the shell edge, caused by one of the unit loads \ud835\udc4b, \ud835\udc4c ,\ud835\udc4d. The edge loads are normal \ud835\udc4b and transverse \ud835\udc4c , while \ud835\udc4d is the moment (Fig. 5). Each of the components in the compatibility equations is related to a certain shell structure, where: \ud835\udc52 \u2014 ellipsoidal shell, \ud835\udc50 \u2014 cylindrical shell. For example \ud835\udeff\ud835\udc522(\ud835\udc4b2) is a vertical displacement component on the second edge (\ud835\udf11 = \ud835\udf112 = \ud835\udf11\ud835\udc522) of the ellipsoidal shell caused by the normal force \ud835\udc4b2 = 1. Obtaining each of the displacement \ud835\udeff and rotation \ud835\udf17 components requires solving Eq. (49) for a specific edge load contained in the vector \ud835\udc53 (Eq. (48)). In the case of \ud835\udeff (\ud835\udc4b2) and \ud835\udf17 (\ud835\udc4b2) the only non-zero \ud835\udc522 \ud835\udc522 loads \ud835\udc4b1\ud835\udc50 , \ud835\udc4c1\ud835\udc50 , \ud835\udc4d1\ud835\udc50 , \ud835\udc4b2\ud835\udc52, \ud835\udc4c2\ud835\udc52, \ud835\udc4d2\ud835\udc52" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001165_s1068371214060042-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001165_s1068371214060042-Figure3-1.png", "caption": "Fig. 3. Characteristics of a WRIG with slip power control.", "texts": [ " Moreover, with this control approach the induction generator operates at its optimum condition. There fore, in addition to higher efficiency and higher power factor, the line current decreases and more power is fed to grid. 2. CONVENTIONAL CONTROL SCHEME In a conventional slip power control scheme (Fig. 2), the speed of the wound rotor induction motor is controlled by varying the external resistance. When, this method is extended for the wound rotor induction generators, the characteristics of the machine change as shown in Fig. 3, which suits a wind turbine based power generation [14]. Figure 3 indicates the different operating points of WRIG under different wind speed. However, in this type of control, there are substantial losses in external resistors at high speed which reduces the overall efficiency of the system [16]. Moreover, at low wind speed, WRIG operate at a low torque condi tion (Fig. 3). Therefore, it this condition, both effi ciency and power factor are low. Thus the performance of WRIG in WECS deteriorates further. 3. PROPOSED SCHEME In the proposed scheme, the characteristics of WRIG are matched with the characteristics of a wind turbine by combined input voltage and slip power con trol (Fig. 4). Here, the slip power controller circuit controls the power in the rotor circuit by varying the external impedance of the rotor circuit. The input voltage is controlled by using an auto transformer (an ac voltage regulator or a tap changing transformer can also be used)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000491_aim.2019.8868892-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000491_aim.2019.8868892-Figure1-1.png", "caption": "Fig. 1. (a) Assistive robot \u201dROSWITHA\u201d (b) Mobile Robot platform \u201dVolksbot\u201d", "texts": [ " Statistics shows that by 2050 the number of people aged over 65 would be doubled [1]. In order to assist the elder people for their daily activities, it is better to have the interactive assistive technology at home, which can obey the command of a user and perform the task as ordered. Therefore, in the Autonomous Systems and Intelligent Sensors Laboratory at the \u201cFrankfurt University of Applied Sciences\u201d, Germany, an assistive intelligent robot called, \u201cROSWITHA\u201d (RObot System WITH Autonomy) [Figure 1a] with different kind of actuators and sensors is being developed. In this paper the \u201dVolksbot\u201d is used as a mobile platform [Figure 1b]. *Supported by Frankfurt University of Applied Sciences, Germany 1S. Sharan, T.Q. Nguyen and P. Nauth are with Faculty of Computer Science and Engineering, Frankfurt University of Applied Sciences, Frankfurt am Main, Germany. s.sharan@fb2.fra-uas.de trungquo@stud.fra-uas.de pnauth@fb2.fra-uas.de 2Rui Arau\u0301jo is with Institute of Systems and Robotics (ISR-UC), and Department of Electrical and Computer Engineering (DEEC-UC), University of Coimbra, Po\u0301lo II, PT-3030-290 Coimbra, Portugal. rui@isr" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002345_s12008-020-00660-1-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002345_s12008-020-00660-1-Figure1-1.png", "caption": "Fig. 1 Main parts of a single-disc clutch system", "texts": [ "eywords Dry friction clutches \u00b7 Thermo-elastic problem \u00b7 Generated frictional heat Indeed, the knowledge of the essential factors that affects the lifetime and behavior of thermo-elastic models for the friction clutch is considered as a crucial issue in industry field. Substantially, the interaction among the essential factors and design parameters for the system of clutch is considered important point for the automobile engineers to select its optimal values of design parameters. Owing to the sliding between the contact elements of a clutch system, the high surface temperatures will be appeared. The main elements (clutch disc, flywheel and pressure plate) of the single-disc clutch are explained in Fig. 1. High-temperature is considered one of the main causes of drawbacks that appear on the contacting surfaces such as cracks, plastic deformations, hot B Oday I. Abdullah odayia2006@yahoo.com 1 Al Furat Al-Awsat Technical University, Najaf, Iraq 2 Department of Energy Engineering, College of Engineering, University of Baghdad, Baghdad, Iraq 3 System Technologies and Engineering Design Methodology, Hamburg University of Technology, Hamburg, Germany spot, etc. Additionally, all the mentioned circumstances may cause the premature damage in the clutch parts" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003279_j.matpr.2020.12.094-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003279_j.matpr.2020.12.094-Figure8-1.png", "caption": "Fig. 8. Total deformations at three distinctive laser power (a) 25 W (b) 50 W (c) 75W.", "texts": [], "surrounding_texts": [ "In this research the preliminary experimentations, RSM-PSO based optimization and thermal analysis followed by stress, strain and deformation results were studied for ES-LBTM approach during the generation of hole profile. It has been observed that for the material ablation characteristics, laser power acts as a dominant parameter out from the other parameters which is studied both from experimentations and simulation results. Moreover, from this research it has also been observed that the laser current intensity promotes laser power to impact a lot during material ablation, with the constant time period and scanning velocity. From the RSM based PSO optimization result it has been obtained the optimal parametric setting of 28 A, 2.6 kHz and 19.28% can give MRR up to 0.0004529 g/s. With this endeavor, an attempt has been made to study the material ablation characteristics in ES-LBTM approach. This research further implied to do more numbers of experimentations followed by analysis of thermal stress and its impacts as it alters depending on the thermal conductivity and thermal expansion coefficient of the material; which in turn widen up the areas in advanced manufacturing approach towards the context of ample range of industrial applications with DSS super alloy. Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper." ] }, { "image_filename": "designv11_22_0002286_j.cagd.2020.101826-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002286_j.cagd.2020.101826-Figure9-1.png", "caption": "Fig. 9. Depth contradiction (Penrose stairs): anamorphic deformation of a discontinuous model S .", "texts": [ " In the example of Fig. 8, we defined each face of the horizontal bar in S with three consecutive planar facets and then bent the central ones. These central facets must be refined, as explained in Section 3.3, to introduce additional degrees of freedom for the desired effect, namely altering the depth ordering between the bar and the central vertical column. The tile texture, aligned with the original edges, help mask the deformation and the resulting shading artifact. In this figure (and also in the upcoming Fig. 9 and Fig. 11), S\u0302 is dilated away from O uniformly, to avoid overlap with S and get a better grasp on the technique. In rigor, the depicted S\u0302 comes thus by composition of a local anamorphic deformation and this global dilation, which results in another anamorphic deformation. Admittedly, we could have constructed this illusion without bending the horizontal bar, by trimming away its central section between the two carefully chosen planes, namely those through O and the diagonally opposed vertical edges of the central column defining its silhouette", " 2) admits a discontinuous polyhedral model S (Ernst, 2006), so the trick to designing its seamless counterpart S\u0302 amounts to closing the gap, by deforming S in an anamorphic manner. The four faces of the vertical column are defined as simple degree-(1,3) B\u00e9zier patches, and then two rows of control points are moved, guaranteeing G1 continuity at the outer faces of the column. The text texture strongly enhances the deception effect, fooling the viewer into thinking that the column is still vertical and made up of planar faces. We could have spread the deformation over all the legs, but, for simplicity, we bent only one. Fig. 9 depicts another example (Penrose stairs), where the overlapping slab requires a local deformation and hence a more elaborate C1 spline representation, with double internal knots. Fig. 10 shows actual pictures of 3D printouts of these models, taken using a white box, confirming that the use of textures results in a convincing effect, even with strong deformations. The models were designed interactively with a graphical model interface in Mathematica\u00ae, which natively supports NURBS and textures in parameter space, and then manufactured with a photo-realistic full-color 3D printer (ProJect CJP 660Pro by 3D Systems)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001428_1350650114559617-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001428_1350650114559617-Figure4-1.png", "caption": "Figure 4. Rollers in the asymmetric state.", "texts": [ " The elastic deformations of the roller between inner and outer ring are i and o respectively and has the relationship R R0 \u00bc i \u00fe o \u00f09\u00de Assuming the roller load is Q, we can obtain i \u00bc 0:2723E0L cos i Dmr 0:074 \" # 1 1:074 Q 1 1:074 \u00f010\u00de o \u00bc 0:2784E0L 1\u00fe o t 0:078 \" # 1 1:078 Q 1 1:078 \u00f011\u00de Therefore, the total elastic deformation of roller with inner ring and outer ring can be expressed as \u00bc i \u00fe o \u00f012\u00de \u00bc 0:2723E0L cos i Dmr 0:074 \" # 1 1:074 Q 1 1:074 \u00fe 0:2784E0L 1\u00fe o t 0:078 \" # 1 1:078 Q 1 1:078 \u00f013\u00de Figure 4 shows that the rolling elements are in the asymmetric state. In this case, the horizontal force is unbalanced if the direction of displacement of inner ring center is still along the loading direction. Therefore, the displacement of the inner ring center cannot move just along the loading direction and the Stribeck method is not suitable here. It is assumed that the angle between roller no. 0 and the external load direction is , the angle between adjacent rollers is \u2019, radial displacement of inner ring center is r, the angle between the direction of radial displacement of inner ring center and load direction is ", " at Purdue University on January 18, 2015pij.sagepub.comDownloaded from where k is the angle between the direction of the displacement of inner ring center and roller no. k, k \u00bc k\u2019 . The loading area is the range of 90 on both sides of on the lower half ring, within this area rollers are loaded, therefore it\u2019s needed to judge the angular position of each roller. The elastic deformation of roller ( k) can be expressed by k \u00bc r cos\u00f0k\u2019 \u00de \u00f015\u00de Otherwise the roller is not loaded, k \u00bc 0. As shown in Figure 4, the roller nos. 0, 1, 2, 7, and 8 are in the loading area. The relationship between their elastic deformation k and the offset parameter of inner center ( r and ) can be determined by equation (15). And the relationship between the elastic deformation of roller and the load is given by equation (13). When r and are given, the load on each roller Qk can be calculated by equations (13) and (15) by Matlab programming. According to force equilibrium and geometric relations, the sum of the components of the load in the vertical direction of each roller is equal to the external load Q0 cos \u00feQ1 cos \u2019 \u00f0 \u00de \u00feQ2 cos 2\u2019 \u00f0 \u00de \u00fe \u00feQk cos k\u2019 \u00f0 \u00de \u00bc Fr \u00f016\u00de The sum of the components of the load in the horizontal direction of each roller should be zero Q0 sin \u00feQ1 sin \u2019 \u00f0 \u00de \u00feQ2 sin 2\u2019 \u00f0 \u00de \u00fe \u00feQk sin k\u2019 \u00f0 \u00de \u00bc 0 \u00f017\u00de Based on force equilibrium equations (16) and (17), we can judge if the offset parameters of inner ring center are correct" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003746_s10409-021-01089-9-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003746_s10409-021-01089-9-Figure15-1.png", "caption": "Fig. 15 Stress distribution of the leaf. a Normal stress xx . b Shear stress xz . c Normal stress zz", "texts": [ " Revolute joints are achieved also by the linear constraints to connect the chassis with the axles and shackles. The spring leaf has a flat bottom and the profile of the top and bottom surface of two wings are described by two cubic polynomials individually. All the geometrical and material parameters are given in Table\u00a01. The weight of the vehicle is assumed to be 2 tones so the mass of the chassis node for one single leaf spring is 500 kg. Figure\u00a013 gives the horizontal and vertical displacement of the chassis node. Figure\u00a014 compares the displacement of two axle nodes at the two direction. Figure\u00a015 presents the 1 3 stress distribution of the leaf at the maximum deformation moment. In this model, a rigid-flexible vehicle suspension model includes the leaf spring, chassis, shackle, axles is efficiently built. The spring eye which consists of steel cylinder, rubber bushing and connects to a rigid axle can be easily modeled by the proposed solid-beam element since it can be assembled at the thickness direction. Although the solid element is also capable, much more degree-of-freedom and integration points should be used" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001829_icra.2016.7487631-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001829_icra.2016.7487631-Figure2-1.png", "caption": "Fig. 2: This figure illustrates the relation between the mobile base and the sensor reference frame. The relative pose of the sensor, with respect to the robot frame, is ks. While the mobile base moves along the trajectory oi, the sensor will move along the path zi in its own reference frame. Lastly, kb represents the robot baseline.", "texts": [ " The path taken by the robot, however has a great influence on the quality of the calibration. Each particular motion along the trajectory might depend only on a subset of parameters. Calibrating the kinematic parameters k of a wheeled mobile platform means estimating the coefficients ko that allow us to compute the relative motion of the base from the ticks measured by the encoders. If the robot is equipped with one or more exteroceptive sensor, these parameters also include the relative position of the sensor ks w.r.t. the mobile base, as shown in Figure 2. Knowing the mobile base parameters ko and the ticks of the wheel encoders ui we can compute the motion of the mobile base during the ith interval, through a direct kinematics function f(ko,ui). Let oT i = (xo i , y o i , \u03b8 o i ) T be this motion. If the robot is equipped with an exteroceptive sensor such as a laser scanner, and it operates on a plane, the sensor parameters are augmented with the sensor position ks = (xs ys \u03b8s)T . Estimating the full kinematics of the robot thus means estimating the vector kT = (koT ksT )" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure15-1.png", "caption": "Fig. 15. Finite element model of internal gear pair with two contact points.", "texts": [ " '' '' '' z x y z t t t t t t t t t t t n n n n l t t \u03b4 \u03b4 \u03b4 \u03b4 \u03d5 \u03d5 \u03b1 \u03b1 \u03d5 \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u0394\u23aa \u23aa \u0394 + \u0394 + \u0394 \u23aa \u23aa + \u23aa \u2202\u03a9 \u2202\u03a9 \u2202\u03a9\u23aa\u0394 = \u00d7 \u22c5 =\u23aa \u2202 \u2202 \u2202\u23a9 \uff08 \uff09 \uff08 \uff09 \uff08 \uff09 \uff08 \uff09 (31) Based on solid modeling methods in Sec. 3, the internal gear models with two contact points can be established as shown in Fig. 14. For its stress analysis, the contact unit size, material properties, boundary conditions and load application are the same with gears with single contact point. The numbers of units and nodes are 378644 and 1608008, respectively. Finite element model of internal gears with two contact points is shown in Fig. 15. The analysis results in Fig. 16 show that the maximum contact stress of internal gears with two contact points is 616.88 MPa. The maximum von Mises stress and shear stress of the pinion with two contact points in Fig. 17 are 417.21 MPa and 133.82 MPa, respectively. The maximum von Mises stress and shear stress of the internal gear with two contact points in Fig. 18 are 347.72 MPa and 165.97 MPa, respectively. Obviously, two peak stress regions can be found on tooth profiles, which are also corresponding to two contact points" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002322_s11665-020-04735-8-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002322_s11665-020-04735-8-Figure12-1.png", "caption": "Fig. 12 Wear mechanism of bionic specimen: (a) early period of wear, (b) subsequent period of wear", "texts": [ " The difference in microhardness among the different samples resulted in different deformation resistance and wear resistance, which further caused differences in furrow depth and density. The increase in micro-hardness can enhance the plastic deformation resistance and is beneficial to improving wear resistance. This result demonstrates that the existence of bionic coupling units was beneficial to improving the wear resistance of 40Cr steel. The change in the wear movement mode reduced the furrow length, indicating that the main wear mechanism was abrasive wear. In the beginning of the wear process, bionic units and substrate were in the same plane (Fig. 12a). When the bionic units moved on the wear pair surface, the wear pair surface near the contacting part produced a certain plastic deformation because of the large pressure between them, which broke the substrate surface near the bionic units. The material produced particles commonly known as wear debris. The grinding debris under the external loads cut into the substrate. Due to the low hardness of the substrate, the cutting depth of the wear debris was large, which plowed out deeper and longer furrows during the wear process. Compared with the hard phase unit, the wear debris was ground off by the unit when it moved to the hard phase bionic unit. Thereby, the cutting depth was reduced and the plow length of the furrow was also decreased during the subsequent wear process. The units could bear the abrasion only until the height difference reached equilibrium position (Fig. 12b). Due to the favorable surface roughness, depth and hardness of the bionic unit, the specimen subjected to LBST exhibited significantly improved wear resistance. Journal of Materials Engineering and Performance In this investigation, the LBST method was demonstrated to be an effective method to improve the wear resistance of brake camshaft with 40Cr steel. The main conclusions of this work can be summarized as follows: 1. In order to obtain smaller surface roughness and larger size of units, the LBST technique was adopted to acquire the appropriate bionic semisolid unit" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003349_s00170-021-06813-0-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003349_s00170-021-06813-0-Figure12-1.png", "caption": "Fig. 12 Ansys Fluent CFD post-processing temperature contour (\u00b0C) showing maximum temperature as 18.64\u00b0C. Active cooling by diaphragm pump and circulatingGalinstan LM coolant through cold plate to reservoir in closed loop. Applied load is 72 W, B.C. is 10\u00b0C from Peltier cooler and 25\u00b0C of natural convection. Presented prototype is in situ AM cold plate with embedded copper electroplated ceramic tubes", "texts": [ " Specific boundary condition available in Ansys (fan boundary condition) is used to split the tube as shown in post-processing (Fig. 11) to apply the characteristics curve of the diaphragm pump presenting flow rate vs pressure for greater fidelity of LM flow model. In solving the energy equation resulting temperature contours, the conjugate heat transfer is used to describe heat transfer that involves variations of temperature within solids and liquid metal due to thermal interaction between the solids and LM as seen in Fig. 12. In this simulation, surface heat is generated from mounted electronics (72 W), and heat is taken away by the AM cold plate and free convection. Heat flux post-processing report has shown that heat transferred by free convection is less than 10% and most of heat dissipated to the cold boundary condition via conduction. CFD simulations performed for in situ AM cold plates including plated ceramic tubes replicating test runs and the pre- and post-processing corresponding to in situ AM cold plate of experiment run #1 (BN-PC-IAM) are shown in Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003180_978-81-322-2352-8_22-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003180_978-81-322-2352-8_22-Figure3-1.png", "caption": "Fig. 3 Nomenclature of the pulse laser sintering process", "texts": [ " The substrate is fixed on the base and the powder was distributed layer by layer on the substrate and the laser beam fell over the surface of powder bed due to which the projected corner of the particles melted and stuck to other particles, forming a block. In this way, the first layer of the powder bed was consolidated. After consolidation of the first layer, the second layer of powder with the same thickness was put over the previously consolidated powder layer and the same process was repeated a number of times. Figure 3 shows the nomenclature of the selective laser sintering. Based on the Taguchi design, L18 orthogonal array was chosen for carrying out the experiment. The controlling parameters, such as, composition, layer thickness, hatching distance, pulse energy, pulse width and powder layer distance from focal plane were the variable parameters. The parameters and their levels are reported in Table 3. Table 4 presents different combinations of parameters for the L18 orthogonal arrays (Phadke 1989; Ross 1996)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001851_s12206-016-0614-7-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001851_s12206-016-0614-7-Figure5-1.png", "caption": "Fig. 5. Illustration of the IMPACT function.", "texts": [ " Therefore, the contact between adjacent leaves is modeled with IMPACT function rather than the \u201ccontact model\u201d. The principle about IMPACT function is mathematically similar to the \u201ccontact model\u201d. The IMPACT function activates when the distance between I and J marker falls below a nominal free length (x1), i.e., when two parts collide. The force is zero when the distance between the I and J markers is greater than x1. An example of a system that can model the IMPACT function is a ball falling toward the ground. Fig. 5 shows the free length value x1 at which the IMPACT force turns on. The force has two components: A spring or stiffness component and a damping or viscous component. The stiffness component is proportional to k and is a function of the penetration of the I marker within the freelength distance from the J marker. The stiffness component opposes the penetration. The damping component of the force is a function of the speed of penetration. The damping opposes the direction of relative motion. The damping coefficient is a cubic step function of the penetration to prevent discontinuity in the damping force at contact" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000384_metroi4.2019.8792915-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000384_metroi4.2019.8792915-Figure4-1.png", "caption": "Figure 4. The solution surface (wheel velocity is normalized)", "texts": [ " Is given by = ( ( ) ) (4) The values of the constants in three typical road conditions are reported in the following Table I: Therefore, the traction curves, in 3 road conditions (tarmac dry and wet, the presence of snow), are given by the following diagram: A typical rule in a Sugeno fuzzy model has the form rules represented in this form (see the following appendix for details): 2 ( ) + ( ) (5) The fuzzy rules for the Tire-Road Interaction are: \u2022 R1: IF is <0.05 OR is low THEN is low \u2022 R2: IF is [0.05;0.15] THEN is medium \u2022 R3: IF is >0.05 OR is high THEN is high The surface result is shown in the following figure: The solution of the T-S model is the surface shown in Fig.4. It is evident that the solution is not univocal but varies with the varying of the different parameters: the surface graphically represents the optimal solution according to the a priori rules (R1, R2 and R3). Obviously, as the rules increase, the complexity of the surface increases. On the other hand, the increase in the variables taken into account the order of the model and therefore it is impossible to represent the solution with a simple three-dimensional surface. IV. THE IOT SURFACE CONTROL SYSTEM In this section we will examine the scheme of the automatic ABS system of a fixed wing drone which is signaled, when it is close to landing, the state of the landing runway surface through a local Wi-Fi: the information will be used to set the right parameters that will then influence the fuzzy logic system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure9-1.png", "caption": "Fig. 9. Backlash caused by center distance change.", "texts": [ " The size of the backlash can be changed by adjusting the length of common normal line or the center distance. As shown in Fig. 8, the backlash as a result of the lesser actual tooth thickness than the ideal non-backlash tooth thickness is given by = \u2212t i ab t t , (44) where ti is the tooth thickness of ideal non-backlash and ta is the actual tooth thickness. With the consideration of eccentricity error, the time variation of the actual geometric center distance of the driven and driving gears will also cause the change in the tooth backlash. As shown in Fig. 9, the change in the backlash caused by center distance change can be expressed as tan\u0394 = \u0394b a \u03c8 , (45) where \u2206a is the change value of geometric center distance. The tooth backlash at any time can be expressed as 2 \u2032 = + \u0394tb b b . (46) Three completely different meshing statuses may exist: (1) non-impact status (non-separation of meshing tooth surface), (2) single-sided impact status (tooth separated, but impact is only generated on the surface of meshing tooth), and (3) double-sided impact status (tooth separated, and impact is generated on both the meshing and non-meshing tooth surfaces at the same time)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003684_j.matpr.2021.05.466-Figure22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003684_j.matpr.2021.05.466-Figure22-1.png", "caption": "Fig. 22. Longitudinal Stress for low carbon steel.", "texts": [], "surrounding_texts": [ "Figs. 16, 17, 18, and 19 shows the result of longitudinal stressinduced in Elliptical, Hemispherical, Torispherical, and Plain formed head. Fig. 23. Circumferential stress for low carbon steel. Fig. 24. Longitudinal Stress for alluminium alloy. Fig. 26. Circumferential stress for gray cast iron. Fig. 27. Longitudinal stress for gray cast iron. Fig. 29. Longitudinal stress for stainless steel. Fig. 30. Longitudinal stress for titanium alloy. Fig. 31. Circumferential stress for titanium alloy. From graph Fig. 20, Stress-induced in Hemispherical head is minimum, while in Plain formed head, stress-induced is maximum. Therefore, for the taken conditions, 8 bar capacity pressure vessels and 24 L, the hemispherical end is better." ] }, { "image_filename": "designv11_22_0003293_j.inffus.2021.01.002-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003293_j.inffus.2021.01.002-Figure4-1.png", "caption": "Fig. 4. The definitions of three coordinate systems in this study.", "texts": [ " In what follows, the proposed framework is described as shown in Fig. 3, which includes coordinate definition, sensor calibration, orientation estimation algorithm, evaluation method of balance ability based on motion capture and computation of spatio-temporal gait parameters, respectively. There are three coordinate systems involved in our measurement system: (1) Navigation coordinate system (NCS); (2) Sensor coordinate system (SCS); (3) Body coordinate system (BCS). NCS is defined as north-east-down coordinate system in Fig. 4. SCS is determined by the sensor itself as shown in Fig. 4. For BCS, the dynamic human model as shown in Fig. 4 is defined. The model takes the pelvis as the root node and traverses up and down to form a complete skeleton kinematic chain. Simultaneously, the kinematic chain is connected by joint points, each of which has its own coordinate system. Setting the shank as example, the axes of \ud835\udc42\ud835\udc65\ud835\udc5d, \ud835\udc42\ud835\udc66\ud835\udc5d and \ud835\udc42\ud835\udc67\ud835\udc5d point the forward, right and down side of the sagittal plane of the human body, respectively. In the field of inertial navigation, the measurement error of sensor itself is a major factor leading to the attitude misalignment [23]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002914_978-3-030-60330-4-Figure3.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002914_978-3-030-60330-4-Figure3.9-1.png", "caption": "Fig. 3.9 Maps of equal frequencies for the rectangular lattice of ellipse-shaped particles (f = 1.5): a LA-mode; b TA-mode; c RO-mode. Curves 1 correspond to = 1.2; 2\u2212 = 1.5; 3\u2212 = 1.9; 4 \u2212 = 2.45; 5 \u2212 = 2.65; 6 \u2212 = 2.95; 7 \u2212 = 3.05", "texts": [ "10 (for f = 1.5). Figure 3.8 shows that anisotropy of the longitudinal mode is most evident when > 2.7, and the longitudinal mode is almost isotropic for < 1.5, whereas the transverse mode is anisotropic even for small . The rotational mode is isotropic for 2.42 < < 2.8. However, all modes are anisotropic even in the low-frequency range in the case of the rectangular lattice with ellipse-shaped particles; see Fig. 3.10. Two modes (LA and RO) are presented at frequency = 3.1 in Fig. 3.10 in addition to Fig. 3.9, where all three modes are separately plotted in the frequency range 1.2 < < 3.05. Here, small \u201crectangles\u201d correspond to the LA-mode and large \u201crectangles\u201d stand for the RO-mode. As it wasmentioned above, the thirdmode (TA) is absent in the system at this frequency. Figure 3.10 shows that the map of equal frequencies reproduces completely the structure of the rectangular lattice at issue. Figure 3.11 contains squares instead of rectangles for the square lattice. (u,w, \u03d5)T = (Au, Aw, A\u03d5)T exp[i(\u03c9t \u2212 k \u00b7 r)], where Au, Aw, and A\u03d5 are the complex amplitudes of harmonic waves, \u03c9 is the oscillation frequency, and k = {kx , ky} is the wave vector, then, according to the procedure described in Sect" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000019_iros.2018.8593927-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000019_iros.2018.8593927-Figure2-1.png", "caption": "Fig. 2: Robot with two Schunk LWA 4P arms in simulation.", "texts": [ " All algorithms are implemented in C++ using the Robot Operating System (ROS), MoveIt! and OMPL. Furthermore, OpenRAVE is used in the implementation of TrajOpt. The evaluation is performed on a standard PC with Ubuntu 16.04 (64-bit), Intel Core i5 CPU with 2.49 GHz and 8 GB memory. The dual-arm robot that is considered as test case consists of two Schunk LWA 4P arms with six degrees of freedom each, which leads to a 12-dimensional configuration space. The arms are mounted on a common basis in a distance of 0.73 m. Figure 2 shows the robot model in the simulation environment. All planning approaches are compared on a set of random motion planning problems. The following procedure has been employed to generate feasible problems for a fixed relative transformation T rel between the end-effectors. Proceeding from a random start state, ngoal = 5 goal states are sampled, whereby a closed chain motion must be possible from start to first goal, from first goal to second goal, and so on. This guarantees that a solution exists between the start state and each goal state" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000566_ab5b65-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000566_ab5b65-Figure9-1.png", "caption": "Figure 9. The cracked tooth on planet gear.", "texts": [ "037 point in order not to lose too much fault information. Derived from equation\u00a0(4), the extensional multiples of Tp , Ts, Tr in MTSA are 3, 10 and 3, respectively. Due to the harsh working conditions and periodic alternating load, tooth crack is a common failure type, which usually occurs at the root of the gear. The development of crack will eventually cause the breakage of teeth or even the entire gear. To detect this fault, the proposed approach is first validated on a planet gear as illustrated in figure\u00a09 where an artificial cracked tooth is introduced to generate fault signatures. Figure 10(a) presents the original raw encoder signal from the setup. Similar to the simulation, it is clearly found that the accumulation of angular displacements is predominant due to its character of monotonically increasing, while the other features are totally indiscernible. To highlight distinguishing details, CDM is implemented to compute the second-order difference signal, which is shown in figure\u00a010(b). However, since the noise is synchronously magnified as well as the meshing variation, slight fluctuations originated from the crack are intractable to be identified from original IAA" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure18.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure18.7-1.png", "caption": "Figure 18.7 Typical investment castings", "texts": [ " Vacuum casting is preferred for casting nickel- and cobalt-base \u2018super-alloys\u2019, not only to avoid contamination of the alloy but also to ensure perfect filling of the mould and to reduce oxidation of the component after casting. When the cast metal has solidified, the mould is broken up to release the castings, Fig.\u00a018.6, which are cut away from their runners. The castings are then finished by fettling, finishing using a moving abrasive belt, or vapour blasting where loose abrasive is forced on to the surface under pressure. Any cores present in the casting are removed by using a caustic solution. Fig.\u00a018.7. The dimensional accuracy of investment casting will vary depending upon pattern contraction, mould expansion and contraction, D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 18 Investment casting, lost foam casting and shell moulding 18 276 XX Virtually unlimited freedom of design enabling intricate shapes, both external and internal, to be produced. Sub-assemblies can often be replaced by a single casting. XX Metals can be cast which are difficult or even impossible to work by other methods" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure4.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure4.5-1.png", "caption": "Fig. 4.5 a Example of an object that should not be grasped with fixed grasping orientation. The object or the gripper would be damaged during the grasp procedure. b Estimation of a plane in the grasp region. c The open vector of the gripper coordinate system has to be aligned with the normal of the plane to ensure safe grasping", "texts": [ "19 Effect of the edge based localization attempt when locating critical planar objects . . . . . . . . . . . . . . . . . . 36 Figure 4.1 Gripper kernel KG used for gripper pose hypotheses generation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 41 Figure 4.2 Pick pose estimation by correlation . . . . . . . . . . . . . . . . . . 42 Figure 4.3 Estimation of the gripper orientation \u03a6. . . . . . . . . . . . . . . . 43 Figure 4.4 Calculation of the approach distance da . . . . . . . . . . . . . . . 43 Figure 4.5 Yaw angle estimation in the context of depth map based bin-picking . . . . . . . . . . . . . . . . . . . . . . . . . . . 46 Figure 4.6 Pitch angle estimation in the context of depth map based bin-picking . . . . . . . . . . . . . . . . . . . . . . . . . . . 46 Figure 4.7 Ambiguities and ambiguity elimination for grasp pose estimation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 51 Figure 4.8 Four example test piles used in the depth map based bin-picking experiments . . . . . " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure3.17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure3.17-1.png", "caption": "Figure 3.17 Squarehead", "texts": [ " The faces are machined to a high degree of accuracy of flatness, squareness, and parallelism, and the 90\u00b0 vee is central with respect to the side faces and parallel to the base and side faces. workpiece square to the reference surface (see Figure 3.8 Wedge Figure 3.9 Jack used to support D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 3 48 3 Marking out This head is used in the same way as an engineer\u2019s square, but, because the rule is adjustable, it is not as accurate. A second face is provided at 45\u00b0 (Fig. 3.17(b)). A spirit level is incorporated which is useful when setting workpieces such as castings level with the reference surface. Turned on end, this head can also be used as a depth gauge (see Fig. 3.17(c)). Figure 3.13 Setting workpiece square to reference surface Figure 3.14 Scribing line square to datum Blade Stock Figure 3.12 Engineer\u2019s square degree of accuracy in straightness, parallelism and squareness. It is available in a variety of blade lengths. The combination set consists of a graduated hardened steel rule on which any of three separate heads \u2013 protractor, square or centre head\u00a0\u2013 can be mounted. The rule has a slot in which each head slides and can be locked at any position along its length" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure18-1.png", "caption": "Fig. 18. Stress analysis results of the internal gear with two contact points (unit: MPa): (a) von Mises stress; (b) shear stress.", "texts": [ " The numbers of units and nodes are 378644 and 1608008, respectively. Finite element model of internal gears with two contact points is shown in Fig. 15. The analysis results in Fig. 16 show that the maximum contact stress of internal gears with two contact points is 616.88 MPa. The maximum von Mises stress and shear stress of the pinion with two contact points in Fig. 17 are 417.21 MPa and 133.82 MPa, respectively. The maximum von Mises stress and shear stress of the internal gear with two contact points in Fig. 18 are 347.72 MPa and 165.97 MPa, respectively. Obviously, two peak stress regions can be found on tooth profiles, which are also corresponding to two contact points. The stress distribution area is relatively concentrated and it has the trends of expanding towards the tooth root direction. The maximum contact stress of tooth profiles with two contact points is 56.8 % lower than that of tooth profiles with single point contact. The maximum von Mises stress and shear stress of the pinion with two contact points are, respectively, 42" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002865_tro.2020.3031885-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002865_tro.2020.3031885-Figure2-1.png", "caption": "Fig. 2. CAD image of the proposed corneal suturing robot.", "texts": [ " The endothelium is the innermost layer of the cornea, which is composed of cells that cannot be regenerated after damage [4]. Therefore, only the top surface and the side cut surface of the cornea are allowed to be touched during the entire surgery. Moreover, suture depth should be thinner than the corneal thickness. In this study, a novel corneal suturing robot is proposed that can produce multiple sutures of the desired shape with high uniformity and accuracy, without touching the endothelium (see Fig. 2). This can be achieved by two major features. First, the robot manipulates the tissue deformation to control the suture shape, instead of manipulating the needle. To manipulate tissue deformation, a novel corneal suturing end effector is proposed. Using the proposed end effector, it can generate multiple sutures with high uniformity and accuracy. Second, a compact origami actuator is attached to the end effector and the needle is inserted into the deformed cornea. The proposed origami actuator is small and able to suture the cornea in every direction without colliding with the surrounding body structure, such as the nose and head bones", " The remainder of this article is organized as follows. In Section II, the structure of the corneal suturing robot and the basic idea of the suturing method is introduced. In Section III, simulation method of determining the end effector shape from the target suture shape is described. In Section IV, the experimental procedure and the method of analysis are explained. In Section V, the simulation results and experimental results are discussed. Section VI concludes this article. The corneal suturing robot is illustrated in Fig. 2(a). The robot is composed of three parts. The first is a five degrees of freedom robotic stage that controls the position and orientation of the end effector regardless of a suture direction. The second is an end effector, a vacuum tweezer (VT), which generates sutures of desired shape with high uniformity. The third is an origami actuator for needle insertion, which is attached on the VT. To overcome the difficulty of multidirectional suturing and to manipulate the VT, five degrees of freedom motion stage is developed Authorized licensed use limited to: Carleton University", " As an XYZ linear stage, the linear piezo stage SLC-2445 from SMARACT is used. The linear stage has a travel range of 29 mm for each axis, 1 nm resolution, and 45 nm repeatability. To fulfill the requirements (R1), (R2), and (R3), a novel suturing module, the VT, is developed. The primary difference between the VT and other suturing methods is that the VT manipulates tissue deformation, not the needle trajectory, to control the suture shape. To manipulate the tissue deformation, the VT has a suction pipe structure with a precisely carved end tip (see Fig. 2). To make a suture, the VT end tip is in contact with the edge of the corneal tissue, and negative pressure is applied to the VT to deform the corneal tissue into a predesigned shape (see Fig. 4). The VT flow line is comprised of three side plates and three frontal plates [see Fig. 4(a)]. The three side plates are equally spaced and parallel to each other. The end-tip of the side plate is in contact with the corneal tissue and determines the deformation shape. The desired deformation can be acquired by appropriately carving the end-tip of the side plate" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001338_s00784-013-1172-3-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001338_s00784-013-1172-3-Figure4-1.png", "caption": "Fig. 4 An example of displacements along the X-, Y-, and Z-axes and brushing force (x10) during a tooth brushing sequence. Each curve on the X-, Y-, and Z-axes and brushing force represents a component cycle of the brushing motion", "texts": [], "surrounding_texts": [ "For each participant, trial, and surface, the program identified no more than 10 cycles that showed the least deviations from the set of four standard scores for each cycle as described below. First, the means and SDs were computed for four measurements (displacement range along the X-, Y-, and Z-axes and the brushing force (BF) range) calculated for each individual from all of their cycles. Each range was defined as the difference between the maximumand theminimumvalues of each displacement andBF within a cycle. Each standard score (SS) was computed by subtracting the individual\u2019s observed cycle value from the mean and dividing the difference by the SD as below: SS X range i\u00f0 \u00de \u00bc Absolute value Mean X range\u2212X range i\u00f0 \u00de=SD X range\u00f0 \u00de SS Y range i\u00f0 \u00de \u00bc Absolute value Mean Y range\u2212Y range i\u00f0 \u00de=SD Y range\u00f0 \u00de SS Z range i\u00f0 \u00de \u00bc Absolute value Mean Z range\u2212Z range i\u00f0 \u00de=SD Z range\u00f0 \u00de SS BF range i\u00f0 \u00de \u00bc Absolute value Mean BF range\u2212BF range i\u00f0 \u00de=SD BF range\u00f0 \u00de This calculated SS indicated a cycle\u2019s amount of deviation from the mean of all cycles for each variable. Overall cycle deviations (OCD) were estimated based on the cycles\u2019 range of displacements and brushing force range using OCD i\u00f0 \u00de \u00bc SS X range i\u00f0 \u00de \u00fe SS Y range i\u00f0 \u00de \u00fe SS Z range i\u00f0 \u00de \u00fe SS BF range i\u00f0 \u00de where SS refers to the standard scores calculated for each individual\u2019s series of tooth brushing cycles. Cycles with the lowest OCD were considered the most representative for each individual surface, for each trial of each participant. It should be emphasized that using only the ten most representative cycles reduces overall variability, and especially betweencycle random variability. This method has been previously applied to analysis of chewing cycle kinematics [13] and has been proven to reduce variation [14]." ] }, { "image_filename": "designv11_22_0003464_lra.2021.3070295-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003464_lra.2021.3070295-Figure1-1.png", "caption": "Fig. 1. Structure diagram of a WMR.", "texts": [ " To describe such an influence, the wheel-terrain interaction involved in slippage is critical for guaranteeing the performance of the WMR teleoperation system. In this section, the kinematic model augmented by slippage was presented to reveal its negative influence on command tracking and system stability. To decrease the command-tracking errors induced by slippage under a velocity-level controller, an acceleration- level controller was adopted, as shown previously [26]. In this study, a two-wheeled differential mobile robot was investigated as a remote robot as shown in Fig. 1, consisting of two active rigid wheels with lugs and one passive wheel. Its kinematic parameters are presented in Table I. In practice, the motion of WMR and wheels can be estimated by position sensors and encoders as Section IV presented. For the WMR above, the ideal assumption of pure rolling is generally existing on hard terrains, which means vi = r\u03c9i (i = 1, 2), and its kinematic model is typically presented as[ vs \u03c9s ] = E [ v1 v2 ] = E [ r\u03c91 r\u03c92 ] = us, (1) where us is the control input and us = [vsd, \u03c9sd]T, and E =[ 1/2 1/2 1/2b \u22121/2b ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003557_s00501-021-01108-z-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003557_s00501-021-01108-z-Figure11-1.png", "caption": "Fig. 11: LMDStrategy (Zone3)", "texts": [ " This involved having two stages, whereby the inclined surface was built first and then partially machined, before deposition on the top of the flash land. This approach resulted in an overall higher quality deposit, since it allowed for the deposition head to remain normal to the die substrate for the majority of the deposition. This was Fig. 12: LMD-pRepair Prior to Machining Fig. 14: RepairedDie inServiceAfter Producing900parts carried out for half of the flash land before being mirrored to the opposite side, as indicated in Fig. 11. With regards to the final machining process, the excess additive material was removed leaving the die within the original tolerances for geometry and surface finish thatwas required by the customer such that their parts were produced without defects. The integration of the LMD-p system at the AFRC was successful and the equipment was able to accurately repair the worn die through the process described. The resulting die was geometrically within tolerance and fit to return to service. The repair at multiple stages is illustrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000654_tnano.2019.2960126-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000654_tnano.2019.2960126-Figure6-1.png", "caption": "Fig. 6. (a) Experimental platform of in vivo evolutionary computation; (b) Microrobot control system; (c) Microchannel device inside the coils.", "texts": [ " The magnetic field used for 2D control can be expressed as: \ud835\udc35 = [ \ud835\udc35\ud835\udc5f sin(\ud835\udf03) cos(\ud835\udf14\ud835\udc61) \ud835\udc35\ud835\udc5f cos(\ud835\udf03) cos(\ud835\udf14\ud835\udc61) \ud835\udc35\ud835\udc5f sin(\ud835\udf14\ud835\udc61) ], (11) where \ud835\udc35\ud835\udc5f , \ud835\udf03, \ud835\udf14, and \ud835\udc61 represent the amplitude of the rotating magnetic field, the direction of motion, the rotating frequency of the field, and run time, respectively. The schematic for Eq. (11) is shown in Fig. 5(c). To validate the computation model, a series of experiments were conducted using a control system that combines the proposed algorithm with feedback control [13]. The control system is composed of a Helmholtz coil system, three power supplies, a National Instruments data acquisition (NI DAQ) device, a camera, a microscope, and a host computer (Fig. 6(a) and (b)). The Helmholtz coil is powered by three power supplies controlled via the DAQ device. Each coil pair generates a magnetic field of approximately 1.5 mT with a current of 0.45 A. The motion of rolling microrobots was recorded using a CMOS camera (BFS-U3-13Y3M-C) at 25 frames per second with a field of view of 614.4 \u00d7 491.5 \ud835\udf07m under a 10 \u00d7 objective lens and a resolution of 0.48 \ud835\udf07m/pixel. The positions of the rolling microrobots were tracked using real-time image processing in LabVIEW. A LabVIEW program was developed to control the strength, frequency, and direction of the magnetic field and the camera, which allows us to implement the feedback control of the rolling microrobots under the proposed algorithm. The microchannel device with the rolling microrobots was installed in the middle of the Helmholtz coil system (Fig. 6(c)). 1536-125X (c) 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. We use both simulation results with MATLAB and experimental results with the microrobot control system to demonstrate the effectiveness of OGSA for in vivo evolutionary computation compared to the brute-force search for the landscapes of Ackley-I and Ackley-II. For the simulation, 12 agents were employed, and the deployment region was set to be \ud835\udc65 \u2208 [\u22125,\u22124], \ud835\udc66 \u2208 [\u22123,\u22122], within which the initial positions of the agents were uniformly generated" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002990_jsen.2020.3042988-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002990_jsen.2020.3042988-Figure10-1.png", "caption": "Fig. 10. (a) Implantable glucose sensor and (b) its components [20].", "texts": [ " To read the fluorescent measurement signal the optical subsystem comprises a 380 nm light-emitting diode (LED as excitation source and two spectrally filtered photodiodes, which measure the fluorescence intensities from the hydrogel and the LED, respectively [20]). The system is remotely powered via a 13.56 MHz inductive link utilizing an extended ISO15693 command set for communication and control of its on-chip analog circuitry to take various optical, thermal, and voltage measurements. The only external support circuitry required is a tuned antenna circuit coupling to the reader field (Fig. 10). This allows the long-term implantation of such sensors below the human skin. The combination of application-specific integrated circuitry, micropackaging and hydrogel-embedded fluorimetry-based sensing has led to the first certified and EU/US marketapproved, implantable glucose sensor solution [48], [20], Eversense by Senseonics [77]. Authorized licensed use limited to: University of Canberra. Downloaded on May 22,2021 at 16:31:09 UTC from IEEE Xplore. Restrictions apply. 1530-437X (c) 2020 IEEE" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000868_gt2015-43971-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000868_gt2015-43971-Figure3-1.png", "caption": "Figure 3: Specimen geometry for stress rupture testing.", "texts": [ "org/about-asme/terms-of-use After laser cladding and/or machining, the surface of the fatigue specimen was initially polished with 400 grit and finally with 600 grit sand paper. All fatigue tests for round IN-625 specimens were performed on a 100 kN Instron 8516 servohydraulic mechanical testing system operated under stress control with sinusoidal waveform and a min/max stress ratio of R=0.1 (tension-tension) and a frequency of 40 Hz. Stress Rupture Test The stress rupture testing was conducted at the bond area between the laser-consolidated IN-738 and cast IN-738 substrate. Figure 3a shows the geometry of stress rupture sample, which is machined from the laser-consolidated coupon as shown in Figure 3b. For comparison, fully LC IN-738 and cast IN-738 stress rupture specimens were also prepared according to Figure 3a. All laser-consolidated and cast IN-738 coupons were heattreated in a vacuum furnace prior to any machining. The stress rupture testing was conducted at 1010\u00b0C (1850\u00b0F) and 55 MPa (8 ksi). LC IN-718 Alloy The LC process successfully produced metallurgically sound IN-718 samples, free of cracks or porosity [10]. SEM observation reveals that the as-consolidated IN-718 has directionally solidified microstructure (Figure 4): the columnar grains growing almost parallel to the build direction, while the horizontal cross section shows fine cell structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003391_10402004.2021.1898708-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003391_10402004.2021.1898708-Figure2-1.png", "caption": "Figure 2. Installation of seal rings: (a) DLC film textured surface SiC seal ring and (b) spiral groove SiC seal ring.", "texts": [ " The structural parameters of the surface texture were set as follows: the texture diameters were 150, 200, and 250 lm, respectively, and the surface densities were 5, 10, and 15%, respectively. Compared to the soft and hard (C-SiC) friction pair, the application of a hard\u2013hard (SiC\u2013SiC) friction pair with high hardness and low expansion coefficients was more favorable to the high-pressure and high-temperature conditions during the working process of the dry gas seal. Hence, in this study the upper and lower specimens were a DLC film textured surface SiC seal ring and a spiral groove SiC seal ring, respectively, as shown in Fig. 2. Friction vibration signal extraction In general, the vibration signal is easily attainable. However, the detected signal contained various interference signals generated by the system, which made the data analysis incapable of reflecting the changes in the state of the seal surface. In this case, it was necessary to obtain the friction-induced vibration signals between seal surfaces by extracting the initial data. In this study, ensemble empirical mode decomposition (EEMD) was performed to extract the friction-induced vibration signals" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001519_s12046-015-0427-x-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001519_s12046-015-0427-x-Figure6-1.png", "caption": "Figure 6. FE model of large diameter bearing (1/2 symmetry model).", "texts": [ " The deviation of experimental results with analytical results is due to the variation in the hardness of the raceways and the ball, the dimensional variations in the ball diameter and deviation in the raceway curvature along the length of the cylindrical groove, the asymmetric in loading resulting in a non-uniform load distribution. Simulation is carried out on an angular contact large diameter ball bearing using ABAQUS Standard FE codes. Abaqus Standard is a general-purpose finite element analyzer that employs implicit integration scheme. Geometrical and physical characteristics of the slewing ball bearing are given in figure 6. For the purpose of study, a 2 m model bearing with 70 rolling elements are modeled. The definition of a nonlinear connector is done in ABAQUS standard FE code. The behavior law is defined along a direction by a load\u2013displacement table. The rings were meshed by \u201cC3D8R: an 8-node linear brick, reduced integration, hourglass control\u201d elements (53,340 elements for each ring) to preserve certain flexibility. The width of these elements at the contact zones, which are tied to the rigid shells, must equal the half-width of the contact ellipse defined by the Hertz theory" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001519_s12046-015-0427-x-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001519_s12046-015-0427-x-Figure3-1.png", "caption": "Figure 3. Bearing ball and race way contact model.", "texts": [ "5.10\u22124 3 \u221a Dw Q2/3, (7) whereas the INA relationship for deformation \u03b4 = 4.84.10\u22124 3 \u221a Dw Q2/3. (8) The above relationships can be used to find the deformation of the rolling bearings. However, because of the highly intimate contact between the rolling element and the tracks and the material nonlinearity, it is most convenient to use the finite element method to determine the nonlinear force-deflection characteristic. 4.1 Numerical determination of stiffness The numerical rolling contact model (figure 3) was used to simulate a contact of the rolling element and a bearing raceway at a varying contact load from 1 kN to 30 kN. The actual model is meshed with \u201cC3D8R: an 8-node linear brick, reduced integration, hourglass control\u201d element in Abaqus standard FE software (Starvin et al 2013). In the direct contact zone, the edges of the contact surface have been divided so as to obtain the dimensions of the segments of the contact surface of about 0.1 \u00d7 0.1 mm. (number of elements on the contact zone: 496, number of nodes in the contact zone: 638)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001392_elektro.2014.6848921-FigureI-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001392_elektro.2014.6848921-FigureI-1.png", "caption": "Fig. I. Schematic representation of static eccentricity", "texts": [ " Eccentricity in Electrical Machines Two different types of eccentricity can be distinguished: \u2022 static eccentricity (SE) \u2022 dynamic eccentricity ( DE) Both of them cause unbalanced magnetic pull (UMP) , i.e. the radial force acting in direction of minimal air-gap. While in case of SE the position of minimal air-gap is constant in time 978-1-4799-3721-9/14/$31.00 mOl41EEE 375 and the value and direction of UMP can be considered invariable in the first approximation, in case of DE the position of minimal air-gap and direction of UMP varies in time. This fact is obvious in Fig. I and Fig. 2. SE is usually characterized by perfectly rigid shaft which is set into the stator with slight misalignment. This state can cause the shaft deflection, particularly in case of long shaft. However this deflection has constant position as stated before. On the other hand, DE is mostly caused by the shaft deflection given during the production or by thermal imbalance. Consequently the shaft oscillation can occur. Besides the shaft deflection or oscillation and the stator core or machine's frame vibration, eccentricity can cause other adverse effects" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003340_tia.2021.3058200-Figure26-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003340_tia.2021.3058200-Figure26-1.png", "caption": "Fig. 26 18/13-pole IFC-DSPMM. (a) Silicon steel sheets. (b) Rotor. (c) Stator. (d) Prototype machine. (e) Test rig.", "texts": [ " The used magnet material is N35H. Because their working temperature may reach 120 \u2103, the demagnetization of the PMs will occur with the circumferential flux density lower than 0.35 T. In the 18/13-pole IFC-DSPMM, the area of the analyzed PM where the circumferential flux density is lower than 0.35 T is much smaller, which indicates that this machine has a better anti-demagnetization capability. An 18/13-pole IFC-DSPMM is manufactured and tested to verify the foregoing 2D FE analysis, as shown in Fig. 26. A stator silicon steel sheet consists of six E-shaped segments. PMs are sandwiched between two adjacent E-shaped segments. As shown in Fig. 26(a), in order to ease assembly difficulties, bridges are introduced to connect the six E-shaped segments, which further aggravates the problem of magnetic flux leakage. In addition, the width of PMs decreases slightly due to the existence of bridges. All of the above factors make the amplitude of the back-EMF 6% smaller than that of the machine without bridges. Meanwhile, it can be seen that there are FC teeth on both sides of each PM to provide magnetic paths. Fig. 26(b)-(c) show the laminated rotor and stator. The prototype machine is shown in Fig. 26(d). PMs are embedded in the adjacent E-type stator core and hence the machine housing is made of aluminum in order to avoid short circuit of magnetic field. Fig. 26(e) is a test rig for torque measurement. Fig. 27 shows that the measured phase back-EMF is basically symmetrical and close to sinusoidal, which agrees well with the FE predictions. Due to end effect, the 3D predicted back-EMF fundamentals are about 9.4% lower than the 2D counterparts. Also, the measured back-EMF fundamentals are about 6.8% lower than the 3D FE predictions owing to manufacturing errors. A method discussed in [23] and [24] is adopted to measure cogging torque. As shown in Fig. 26(e), the stator is clamped by the jaws of the lathe, while the rotor shaft is connected to a balanced beam whose one end has a bolt inserted vertically. There is a metal block welded to each side of the bolt. The lower metal block is supported by the digital gauge, while the upper one can be rotated to adjust the balanced beam horizontally. When the balanced beam has been leveled, i.e., the value of the gradienter is zero, the digital gauge resets to zero. By rotating the indexing head, the stator rotates accordingly and thus the net mass acting on the digital gauge can be measured at different relative positions between the stator and rotor, eventually calculating the torque acting on the rotor shaft" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002355_s10846-019-01096-w-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002355_s10846-019-01096-w-Figure3-1.png", "caption": "Fig. 3 Match based on the sum of paired wheels or difference between the paired wheels\u2019 L2 distance", "texts": [ "3Wheel-Matching Procedure Our local wheel detector does not distinguish between the four wheels of the wall-climbing robot (in fact, these are difficult to distinguish), so we need to match its wheels between video frames to ensure consistent calculation of the robot\u2019s direction angle. For four wheels Wi, i \u2208 {1, 2, 3, 4} detected in frame I and four wheels Wj, j \u2208 {1, 2, 3, 4} detected in frame J, we need to correctly match the wheels Wi and Wj . We find that matching cost functions with the sum \u2211 i,j Dij or the difference (max(Dij ) \u2212 min(Dij )) in the L2 pixel distances between matched wheels both cause mismatches, as show in Fig. 3. We therefore propose the matching cost function shown in Eq. 2, which combines these two cost functions: cost = \u2211 i,j Dij + \u03b3 \u2217 (max(Dij ) \u2212 min(Dij )) (2) where Dij represents the L2 distance of two paired wheels, and \u03b3 is used to adjust the weight of (max(Dij )\u2212min(Dij )) in the whole cost function. We tested the proposed wheel matching cost function by simulation and experiment. In the simulation test, we set the size of the wall-climbing robot to 40\u00d750 cm, and the initial posture of the robot [x, y, \u03b8 ] to [0, 0, 0]", " When performing wheel matching, we match the four wheels of the wall-climbing robot at time T with the wheels at time T \u2212 Nf \u2217 dt , where Nf is the number of time steps (the greater the Nf , the larger distance the wheel will move). Figure 5 shows the curve between the number of failed matches and the value of Nf when \u03b3 takes different values. As can be seen from Fig. 5, when \u03b3 = 0, the number of failed matches increases as Nf increases (when Nf increases, the moving distance of the wheel increases, and the failed match shown in Fig. 3a will occur). When \u03b3 > 0, the number of failed matches hardly changes as Nf increases. The simulation results show that our proposed wheel matching cost function (\u03b3 > 0) can significantly reduce the number of failed matches when the move distance of wheel is large. In the Appendix Videos of video 2.mp4 and video 3.mp4, we used \u03b3 = 0.4 when performing wheel matching. As can be seen from the videos, the wheels are always correctly matched during the movement of the wall climbing robot. The task of the global bounding box detector is to locate the position of the wall-climbing robot in the entire video frame (thus determining the position of the ROI box)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003430_09544070211004507-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003430_09544070211004507-Figure2-1.png", "caption": "Figure 2. Structure of rigid ring tire and PM BLDC motor.", "texts": [ "27 The belt ring unit includes the belt, the tread, and a part of the sidewall. The sidewall is simplified as wheels-ring springs and dampers in the radical and rotational directions. The grounding springs are introduced to express the deformation of the tire and calculate the contact force in the vertical direction between tire and ground. The pure longitudinal sliding brush model is introduced to calculate the tangential force in the imprint. The structure of the rigid ring tire model combined with PM BLDC model is shown in Figure 2. where kt_rt and ct_rt are rotational stiffness and damping coefficient of tire sidewall, kt_rd and ct_rd are radial stiffness and damping coefficient of tire sidewall. mms, mt are mass of stator and tire ring respectively, mw_mr is the mass of the wheel and rotor. FUMPz and FUMPx are UMP in vertical and longitudinal direction respectively. kbeaz and kbeax are motor bearing stiffness in vertical and longitudinal direction respectively. kcz is residual stiffness. kcp is tread element horizontal stiffness" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003519_s10846-021-01359-5-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003519_s10846-021-01359-5-Figure2-1.png", "caption": "Fig. 2 Wheel configuration and parameters of the Four Mecanum Wheeled Robot", "texts": [ " The methodology is based on a scenario where person \u201cA\u201d guides a person \u201cB\u201d by holding his hand, wherein person \u201cB\u201d moves in proportion to the magnitude of the pull towards the direction indicated by person \u201cA\u201d. Using this, the current study attempts to put forth a collaborative method to carry and transport objects using MeWBots. The MeWBot equipped with Mecanum wheels has a passive roller placed at an angle of 45\u25e6 with respect to the wheel axis. When each of the wheels is actuated, these rollers tend to provide a reaction force in a direction normal to the roller axis, as shown in Fig. 2 with green vector arrows. The summation of these forces produced by each wheel provides the robots with an overall direction of motion. Consider the following notation for gaining a better understanding: \u2013 YGX as the global coordinate frame. \u2013 YROXR as the robot\u2019s coordinate frame. \u2013 (\u03c91, \u03c92, \u03c93, \u03c94) be respective wheel velocities in rads\u22121. \u2013 R be the radius of wheel in meters. \u2013 \u03c9Z be the angular velocity about an axis perpendicular to both XR and YR in rads\u22121. \u2013 Vx be the translation velocity of the robot in XR direction in ms\u22121" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002767_jsen.2020.3025054-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002767_jsen.2020.3025054-Figure2-1.png", "caption": "Fig. 2. An illustration of the UAV\u2019s movement.", "texts": [ " When the UAV flies towards the PT, the sensing result will be more accurate due to the fact that the sensing SNR is higher. Then, the UAV will have more opportunities to utilize the licensed spectrum and the throughput of the UAV system can be improved. However, the propulsion energy due to UAV\u2019s movement increases with the flying distance. Hence there are some tradeoffs between the throughput, the EE and the flying distance. In this work, the placement of the UAV will be optimized to satisfy different requirements of the UAV system. An illustration of the UAV\u2019s movement is shown in Fig. 2. It is assumed that the final position of the UAV is Dfm away from the PT. Since the UAV do not have the knowledge of the PRs\u2019 locations, it needs to consider the worst case locations of the PRs. If Df > rp, the UAV should assume that the PR is located on the protected boundary nearest to it, i.e., the distance between the UAV and PR is lpp =\u221a H2 + (Df \u2212 rp)2; if Df \u2264 rp, the PR is assumed to have the same horizontal location with the UAV, in this case, the distance between the UAV and PR is H. According to [2], the hovering power of the UAV is Ph = [ (mug) 3 2\u03c0d2uKu\u03c9 ] 1 2 , (1) where mu, \u03c9 and g are the mass of the UAV in (kg), air density Authorized licensed use limited to: University of Saskatchewan" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002854_j.engfailanal.2020.105028-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002854_j.engfailanal.2020.105028-Figure1-1.png", "caption": "Fig. 1. 3D and FE model of the connecting rod assembly.", "texts": [ " This paper takes the aluminum alloy connecting rod of aero kerosene piston engine as the research object, analyzes the constraint relationship between connecting rod components, determines the boundary conditions of connecting rod on aviation kerosene piston engine, studies the safety factor of connecting rod under thermal mechanical coupling conditions. Finally, a new design is proposed to improve the fatigue strength of the connecting rod. In order to analyze the more realistic working condition, the connecting rod assembly is divided by the second-order tetrahedron element and the first-order hexahedron. The composition and meshing of the connecting rod are shown in Fig. 1 and the meshing results are shown in table 1. Some lubricating oil passage and fillet inside the connecting rod are processed in details. Important joins between components were meshed with matching mesh patterns on mating surfaces. Fillet is not good for meshing, which will cause difficulties in structured mesh generation, leading to smaller mesh size and worse quality. In order to ensure the mesh quality and calculation accuracy, the fillet surface is extracted and meshed. The mesh quality before and after extraction is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure1-1.png", "caption": "Fig. 1. Geometric model of helical gear with geometric eccentricity.", "texts": [ " Finally, the accuracy of the results is verified by comparison between simulation and experiments. The rotation centers of driving and driven gears do not coincide with the geometric centers because of the existence of eccentricity error. e1 and e2 are the eccentricity of driving and driven gears, respectively. The driving gear is rotating anticlockwise. The origin O of absolute coordinate system OXYZ is located at the center O1 of the driving gear rotating center, and O2 is the rotating center of the driven gear, as shown in Fig. 1. O1\u00b4 (x1, y1) and O2\u00b4 (x2, y2) are the geometric centers of driving and driven gears, respectively. \u03a8 is the complementary angle of between line of centers and the meshing line, which varies over time t. The meshing line of the gear pair is tangent to the base circle at points A and B. The meshing line and lines O1O2 intersect at point P. \u03b1\u0384t is the angle between the meshing line and axis Y, which is also the dynamic meshing angle of the gear pair. At any time t, O1\u00b4 (x1, y1) and O2\u00b4 (x2, y2) can be expressed as 1 1 1 1 1 1 1 1 2 2 2 20 2 2 2 20 cos( ) sin( ) cos( d ) sin( d ) ", " 1 b b b b kx x r k ry y k kx x r k ry y k = + + = \u2212 + = \u2212 + = + + (9) The equation of the dynamic meshing line through two tan- gent points A and B is shown as 02 02 01 02 02 01 \u2212 \u2212= \u2212 \u2212 y y y y x x x x . (10) The abscissas of point P can be obtained by letting y = 0, that is, 02 01 02 02 02 01 \u2212= \u2212 \u22c5 \u2212P x xx x y y y . (11) Then, the instantaneous transmission ratio of the gear pair with eccentricity can be described as 1 12 2 \u2212= = P p H xi x \u03c9 \u03c9 , (12) where the given input of \u03c91 is a constant value, and the timevarying \u03c92 can be obtained by combining Eqs. (1)-(12). As shown in Fig. 1, v(t) is the linear velocity of the gear in the direction of the meshing line at any time t. 1( ) cos .p tv t x\u03c9 \u03b1\u2032= (13) In a meshing period, the change diagram of the ideal errorfree helical gear contact line length is shown in Fig. 2. b is the meshing tooth width, LCD is the actual meshing line length of the gear pair end face, L is the length of contact line, and \u03b2b is the helix angle of base circle. According to the relationship between the transverse contact ratio \u03b5\u03b1 and axial contact ratio \u03b5\u03b2, helical gears can be divided into two cases, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000802_978-3-319-13963-0-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000802_978-3-319-13963-0-Figure16-1.png", "caption": "Fig. 16. The geometry of the improved manipulator and two details of CAD prototype", "texts": [], "surrounding_texts": [ "In this paper, an error model of a 2-DOF high-speed parallel manipulator is presented and simulated numerically. The influence coefficient is defined so that the effect separate geometric errors have on the pose errors of the end effector can be better compared. This helps guide the manufacture and assembly of prototypes. Among the parameter errors, the length errors of the distal links and the length errors of the connecting bars have the greatest effect on the pose error. Based on the error model developed, the distribution of pose errors in the workspace reveals the relationship between input parameter errors and output pose errors and provides theoretical support for calibration of the manipulator. In addition, the structure of an improved manipulator combining the 2-DOF manipulator with a rotating mechanism is developed. The dynamic character of the manipulator is yet unknown and needs further research to demonstrate the potential application value of the improved manipulator. Acknowledgments. This research was supported by the National Natural Science Foundation of China (Grant No.91223201), the Natural Science Foundation of Guangdong Province (Grant No.S2013030013355), Project GDUPS (2010) and the Fundamental Research Funds for the Central Universities (2012ZP0004). These supports are greatly acknowledged. 110 J. Wei et al." ] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure16.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure16.6-1.png", "caption": "Figure 16.6 Four-high mill", "texts": [ "4 Rolls for producing a tee section Finished section First pass other. A platform is positioned such that it can be raised to feed the metal through in one direction or support it coming out from between the rolls in the opposite direction or be lowered to feed the metal back through or to support it coming out from between the rolls in the opposite directions. In cold rolling, the roll pressures are much greater than in hot rolling \u2013 due to the greater resistance of cold metal to reduction. In this case it is usual to use a four-high mill, Fig.\u00a016.6. In this arrangement, two outer rolls of large diameter are used as back-up rolls to support the smaller working rolls and prevent deflection. and between the centre and lower rolls in the When rolling strip, a series of rolls are arranged in line and the strip is produced continuously, being reduced by each set of rolls as it passes through before being wound on to a coil at the end when it reaches its final thickness. Extrusion usually has to be a hot-working process, due to the very large reduction which takes place during the forming process" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001906_acc.2016.7525490-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001906_acc.2016.7525490-Figure1-1.png", "caption": "Fig. 1. Quanser 3 DOF Hover.", "texts": [ " For full-state feedback, the FPRE control law is given by u(t) = K(t)x(t), (3) where the feedback gain K(t) is given by K(t) = \u2212R\u221212 BT(x(t))P (t)x(t), (4) and P (t) is the solution of the forward-in-time Riccati equation P\u0307 (t) = AT(x(t))P (t) + P (t)A(x(t)) \u2212 P (t)B(x(t))R\u221212 BT(x(t))P (t) +R1, P (0) \u2265 0, (5) where R1 \u2208 Rn\u00d7n is positive semidefinite and R2 \u2208 Rn\u00d7m is positive definite. The weighting matrices R1 and R2 can also be state-dependent, i.e. R1(x(t)), R2(x(t)), which introduces additional degrees of freedom in the controller synthesis. III. 3DH TESTBED The 3DH testbed [17] shown in Fig. 1 consists of a planar round frame with four propellers. The frame is mounted on a joint that enables the body to rotate about three axes. The propellers are driven by four DC motors. The lift force generated by the propellers is used to control the pitch and roll angles. Yaw control is done using the total torque generated by the propellers. The aim is to control the pitch and roll of the 3DH while maintaining constant yaw. The nonlinear equations of motion of 3DH are given by \u03c6\u0307 = p+ sin\u03c6 tan \u03b8 q + cos\u03c6 tan \u03b8 r, (6) \u03b8\u0307 = cos\u03c6 q \u2212 sin\u03c6 r, (7) \u03c8\u0307 = cos\u03c6 cos \u03b8 r + sin\u03c6 cos \u03b8 q, (8) p\u0307 = Jyy \u2212 Jzz Jxx qr + 1 Jxx \u03c4r, (9) q\u0307 = Jzz \u2212 Jxx Jyy pr + 1 Jyy + 1 Jyy \u03c4p, (10) r\u0307 = Jxx \u2212 Jyy Jzz pq + 1 Jzz \u03c4y, (11) where \u03c6, \u03b8, \u03c8 are the Euler angles, p, q, r are the angular rates in the body axes, Jxx, Jyy , Jzz are the moments of inertia, and \u03c4r, \u03c4p, \u03c4y are the torques acting on the roll, pitch, and yaw axes, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002274_j.mechmachtheory.2019.103748-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002274_j.mechmachtheory.2019.103748-Figure3-1.png", "caption": "Fig. 3. Cylindrical contact and normal reaction distribution.", "texts": [ " A common simplifying factor is the relatively small dimensions of the contact area. Another possibility is to dismember a single coupling into two or more couplings acting in parallel (e.g., see Fig. 4 in Section 3.2 ). The cross section geometry of the slider shown in Figs. 1 and 2 also influences the normal reaction distribution and, consequently, the normal reaction and friction force magnitudes. In general, the coefficient of friction available in the literature is for planar surfaces only. In Fig. 3 , the normal reaction distribution for an external applied force Q that passes through a cylindrical axis is illustrated. The angle of engagement of the pin with the insert is denoted by 2 \u03b2 . The common approach is to adjust the coefficient of friction according to the normal reaction distribution law. For cylindrical surface, Baranov [19] recommends the use of two laws: constant, to be used for new pairs; and cosine, to be used for old pairs that have worn and accommodated. In this paper, it is assumed full contact, 2 \u03b2 = \u03c0rad, old pairs; therefore, the appropriate coefficient of friction is obtained multiplying the coefficient of friction for the planar case by 4 \u03c0 when the contact surface is cylindrical" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002523_j.matpr.2020.05.322-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002523_j.matpr.2020.05.322-Figure2-1.png", "caption": "Fig. 2. Modelling of anti-roll bar bend.", "texts": [ " Naturally, six degrees of opportunity happens at every hub. This component is s of Anti Roll Bar. at a deformed Scale. ng and finite element analysis of anti-roll bar using ANSYS software, Mate- Fig. 9. X Axis Static Nodal Stress. appropriate for direct, enormous revolution, and huge strain nonlinear applications. BEAM189 can be utilized with any bar crosssegment characterized in the program. Versatility, creep, and pliancy models are supported. BEAM189 overlooks any genuine consistent information. Fig. 2 shows modelling of anti-roll bar bend [17,18]. Please cite this article as: V. Mohanavel, R. Iyankumar, M. Sundar et al., Modelli rials Today: Proceedings, https://doi.org/10.1016/j.matpr.2020.05.322 COMBIN14 is a spring-damper mix component that has longitudinal or torsional ability in one, two, or three dimensional applications. interpretations in the nodal , y, and z headings. Revolutions about the nodal , y, and z tomahawks. No twisting or pivotal bur- c Nodal Stress. ng and finite element analysis of anti-roll bar using ANSYS software, Mate- Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003440_j.mechmachtheory.2021.104297-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003440_j.mechmachtheory.2021.104297-Figure2-1.png", "caption": "Fig. 2. Kinetostatic parameters of the GSP limbs.", "texts": [ " In this model, the limb is considered as a chain composing three linear elastic elements: an actuator of the translational motion, and two spherical joints connecting the limb with the moving platform and the base. Since the GSP presents a parallel-kinematics system, its stiffness matrix is expressed through equation K = J T K limb J (67) where K limb is the diagonal matrix combining stiffness values k i of the limbs ( i = 1, 2, \u2026, 6), and J is the Jacobian matrix, K lumb = Diag [ k 1 , k 2 , . . . , k 6 ] (68) J = \u23a1 \u23a2 \u23a2 \u23a3 e T 1 n T 1 e T 2 n T 2 ..... ...... e T 6 n T 6 \u23a4 \u23a5 \u23a5 \u23a6 (69) Kinetostatic parameters of the GSP limb are shown in Fig. 2: e i ( i = 1, 2, \u2026, 6) is the unit vector of the i th limb orienta- tion, and n i is the moment of the unit vector e i with respect to the center of the moving platform O p , e i = r i / l i and n i = b i \u00d7 e i , i = 1 , 2 , . . . , 6 (70) where r i is the vector of the limb M i Q i ( i = 1, 2, \u2026, 6); b i is the vector of the i th limb arm relative to point O p ; and l i is the length of the i th limb, r i \u2261 [ x i y i z i ] = r Qi \u2212 r Mi = [ x Qi \u2212 x Mi y Qi \u2212 y Mi z Qi \u2212 z Mi ] = [ r 2 cos \u03b2i \u2212 r 1 cos \u03b1i r 2 sin \u03b2i \u2212 r 1 sin \u03b1i z i ] (71) l i = \u2016 r i \u2016 = \u221a x 2 i + y 2 i + z 2 i (72) \u03b1i and \u03b2 i are the angles defining the point of the i th limb applications in the base and moving platform, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003627_icedme52809.2021.00056-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003627_icedme52809.2021.00056-Figure2-1.png", "caption": "Figure 2. ROPE RIDE wall-climbing robot", "texts": [ " One is used to generate upward thrust and the other is used to generate thrust in the vertical direction to make it close to the wall. The ability to overcome obstacles is strong, and the requirements for the wall are not high. However, affected by the external wind, the wall-climbing robot occupies a large space, which limits the form of operation. In 1995, Lecturer Xiliang from Osaka Prefecture University in Japan took the lead in researching such a thrust-adsorbing wall robot[5]. Later, in 2013, Seoul University in South Korea developed ROPE RIDE [6]. As shown in figure 2, the crawler movement method is adopted. Under the condition of a load of 20 kg, the climbing speed can reach 15m/min. China's research on thrustadsorption wall-climbing robots is relatively late, but it has developed rapidly in recent years. Yang Meiqiang of Beijing Institute of Petrochemical Technology developed an underwater thrust-adsorption wall-climbing robot prototype in 2015, which is mainly used for observation and detection of underwater structures. 1) Crawler type permanent magnet adsorption wallclimbing robot Permanent magnet adsorption is simpler than electromagnetic adsorption" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000549_0731684419888588-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000549_0731684419888588-Figure5-1.png", "caption": "Figure 5. Analytical model of 3D CTCS under in-plane compression. (a) 3D cell, (b) 2D cell, (c) twist angle, and (d) cell wall.", "texts": [ " At the same time, since the deformation of the 2D cell is not symmetric, the whole structure has a twist angle /, which presents the compression-twist effect. Therefore, the elastic properties of this 3D cellular structure can be determined by effective Young\u2019s modulus Ee and relative twist angle /R, which are written as follows Ee \u00bc r ey (4) /R \u00bc / ey (5) t \u00bc ex ey (6) where r is the stress of cellular structure. Analytical models of the 3D cell under in-plane compression To analyze the compression-twist effect, the mechanical analytical model of the 3D cell under in-plane compression is carried out, which is shown in Figure 5. In the mechanical analytical model, since the bending deformation of the cell wall is relatively small, the geometric nonlinearity of the cell wall is not considered, and the influence of the axial and shear deformation on the deflection of the beam is also neglected. In addition, according to the results of FEA in the following section, the deformation of the ring can be also neglected. Therefore, in the process of the calculation, the ring in the 3D cell could be regarded as a rigid structure. The compressive loading is shown in Figure 5(a), and the number of cell vertex is from A to H. Since the twist deformation occurs mainly in the 2D cells of the vertical plane, the bending deformation of the cell wall is symmetric in every vertical plane. Therefore, the vertex B is chosen as the object, and the B and B1 are different vertexes in the XY plane and YZ plane. Assuming that the unit cell is subjected to a resultant force P in the y direction, the deformation diagrams of ligament B and ligament B1 are shown in Figure 5(d), and their deflection can be written as follows d \u00bc Psinolc 3 3EI Tlc 2 2EI (7) where I is the second moment of inertia of the cell wall, E is Young\u2019s modulus of cell\u2019s material, lc \u00bc ffiffi 2 p 2 lcosh is the length of ligament, o\u00bc 458 h is the angle between ligament and symmetry axis, and T is the applied moment, which can be written as follows T \u00bc Plcsino 2 (8) where P is a resultant force applied on ligament P \u00bc rtl (9) where r is the compressive stress of cell; therefore, the deflections of B and B1 can be written as follows dB1 \u00bc ffiffiffi 2 p rtl4sin 45 h\u00f0 \u00decos3h 48EI (10) dB \u00bc ffiffiffi 2 p rtl4sin 45 \u00fe h\u00f0 \u00decos3h 48EI (11) When, the I \u00bc t4 12 and t \u00bc al, the equations (9) and (10) can be written as follows dB1 \u00bc ffiffiffi 2 p rlsin 45 h\u00f0 \u00decos3h 4Ea3 (12) dB \u00bc ffiffiffi 2 p rlsin 45 \u00fe h\u00f0 \u00decos3h 4Ea3 (13) Therefore, the strain of cell in the y direction is as follows ey \u00bc dB1sin 45 h\u00f0 \u00de \u00fe dBcos 45 h\u00f0 \u00de l (14) The effective Young\u2019s modulus can be written as follows Ee \u00bc 2 ffiffiffi 2 p Ea3 cos3h (15) In the vertical plane, the deflections of B and B1 in the x and z directions can be written as follows dBx \u00bc dBsin 45 h\u00f0 \u00de (16) dB1z \u00bc dB1cos 45 h\u00f0 \u00de (17) When the deformation is small, the twist angle is between the undeformed ligament and the deformed ligament, which is shown in Figure 5(c). Therefore, the relative twist angle that is the twist angle per strain is as follows /R \u00bc arctan dB1z\u00fedBx l dBx\u00fedB1z ey (18) Combining above equations gives /R \u00bc arctan ffiffi 2 p rcos3hcos2h 4Ea3ffiffi 2 p rcos3h 4Ea3 (19) The strain of cell in the x direction is as follows ex \u00bc dBxsin 45 h\u00f0 \u00de\u00fedB1zcos 45 h\u00f0 \u00de l (20) Therefore, the Poisson\u2019s ratio of the cell is as follows t \u00bc cos2h (21) In the above analytical model, when the cell is applied by the compressive loading, it has compressive deformation in the vertical direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001656_tmag.2014.2330812-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001656_tmag.2014.2330812-Figure6-1.png", "caption": "Fig. 6. (a) SPRMM rolling along a 2D spiral moving path in a clockwise direction. The blue square indicates the processed image of the SPRMM. The green cross shows the central pixel coordinate of the SPRMM. (b) Rolling trajectory of the SPRMM along the spiral moving path.", "texts": [ " The magnetization of the NdFeB magnet is 9 55 000 A/m. To make the SPRMM continuously roll along a surface without slip, M1 and M2 were built with the geometric parameters shown Fig. 5 [5], and the outer surface of M2 was painted blue so that it could be effectively tracked by blue color image tracking. We first observed the rolling locomotion of the SPRMM along a predefined spiral rolling path on a horizontal thin surface using open-loop control and the proposed closed-loop control methods, as shown in Fig. 6. Sequential spiral moving paths of the SPRMM were manually set with 150 sequential target points using a graphic interface in the control panel. We set the rotating speed and magnitude of the ERMF as 0.5 Hz and 14 mT, respectively, which are sufficiently strong to control the SPRMM [5]. To generate the rolling locomotion of the SPRMM along the spiral moving path with the openloop control method, sequential ERMFs varying with respect to time were precalculated to the sequential target points, as shown in Fig. 3(c), using (3) and (4). The ERMFs were sequentially applied to the SPRMM without consideration of TABLE I MAJOR SPECIFICATIONS OF THE MNS Fig. 7. (a) and (b) xy plane and xz plane images of the predefined moving path of the SPRMM on a 3D arbitrary surface captured by the biplane camera. (c) Rolling trajectory of the SPRMM. the position of the SPRMM. Fig. 6(b) shows that the SPRMM increasingly deviates from the predefined rolling path as it moves to the final target point. The maximum position error was measured to be around 10 mm. The SPRMM was more likely to deviate as the rotating speed of the ERMF increased due to inertial effect [5], [7]. On the other hand, the SPRMM controlled by the proposed closed-loop control method showed fine, rolling locomotion along the predefined moving path and position errors of less than 1.8 mm, as shown in Fig. 6. The maximum allowable position (distance) error at every target point was set to 2 mm. The SPRMM showed fine, controlled motion with rotating speeds ranging from 0.1 to 5 Hz. We then observed the complex rolling locomotion of the SPRMM along a 3D rolling path on an arbitrary thin surface, as shown in Fig. 7. In the open-loop control of the SPRMM on the complex 3D surface, as shown in Fig. 7, the SPRMM was deviated or got lost from the predefined moving path. However, in the closed-loop control, the SPRMM showed fine, controlled rolling locomotion along the predefined rolling path except three large position errors" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001545_1350650115602510-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001545_1350650115602510-Figure1-1.png", "caption": "Figure 1. Journal bearing scheme.", "texts": [ " This model allows to obtain an analytical description of the fluid film force and of the stiffness and damping coefficients for several values of the Reynolds number and of the L/D ratio in the case of finite bearing. Theoretical analysis The scheme here analysed consists of symmetric balanced rigid rotor on two equal cylindrical full journal bearings. The symmetry about the rotor middle plane allows study one of the parts into which the system is subdivided by the above mentioned plane. With reference to the journal bearing scheme in Figure 1, under the hypothesis of isothermal hydrodynamic lubrication, the equation which governs the pressure in the lubricant film is in non-dimensional form: 1 R2 @ @ h3 12 kx @ p @ ! \u00fe @ @ z h3 12 kz @ p @ z ! \u00bc 1 2 ! @ h @ \u00fe 2 @ h @t \u00f01\u00de where kx and kz are functions for taking in account the turbulence effects proposed by Constantinescu6 and based on the Prandtl\u2019s mixing length hypothesis; h \u00bc C 1\u00fe \"cos \u00f0 \u00de is the film thickness, t the time and the dynamic viscosity of the lubricant. The journal centre position is determined in the rotating frame by the polar coordinates C\", \u2019\u00f0 \u00de", " In Figures 2 and 3 are reported the plots of the steady fluid film components fr and ft, with respect to the eccentricity ratio for several values of Rm. Rotor dynamic behaviour In the case of a rigid, symmetrical rotor with a mass of 2m, spins with constant angular velocity around its axis, and supported in two similar, plain, cylindrical short journal bearings, the rotor is allowed two degrees of freedom such that every point of the rotor performs the same planar motion. By using the Cartesian coordinates system (Figure 1) describing the motion of the journal centre, the dimensionless motion equations may be written as m \u20acx \u00bc fx \u00bc fx x, y, _x, _y\u00f0 \u00de m \u20acy \u00bc fy \u00bc fy x, y, _x, _y\u00f0 \u00de \u00f06\u00de with x \u00bc x C y \u00bc y C m \u00bc CM!2 W at UNIV CALIFORNIA SAN DIEGO on December 24, 2015pij.sagepub.comDownloaded from With the transformation to fixed system coordinates (Figure 1) the non-dimensional fluid film force components fx and fy can be calculated. fx fy \u00bc sin \u2019 cos \u2019 cos \u2019 sin \u2019 fr ft The analytical expressions of fluid film force make possible to obtain easily the stable trajectories of the centre of the journal by integrating numerically by using a Runge\u2013Kutta numerical algorithm. As an example in Figure 4 are reported the trajectories of the journal centre in correspondence of a set of typical stable behaviour of the system, by varying only the values of t Rm (Rm\u00bc 5000; Rm\u00bc 10,000; Re\u00bc 15,000)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001992_icma.2016.7558710-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001992_icma.2016.7558710-Figure4-1.png", "caption": "Fig. 4. Model of the general rigid-body biped robot configuration.", "texts": [ " There are a few studies with regard to toe and heel clearance to the ground during swing phase [14], [21], and we focus on the heel clearance which has simpler pattern than toe clearance. The heel clearance starts increasing at tHR in single support phase and has maximum at tFA in swing phase. After that, it decreases until the heel contact to the ground at the end of stride cycle (tNI ). We fix the maximum of the heel clearance at 15% of the stature by referring to the studies. We can generate natural walking pattern for a general rigid-body biped robot configuration described in Fig. 4, by incorporating normal walking characteristics described in the previous section. We first find pelvis trajectory and the foot orientation with respect to the ground. Then we can solve inverse kinematics during support phase. For swing phase, we obtain initial and final conditions from the support phase and use some constraints which we define for generating the ankle trajectory. After the ankle trajectory is determined, we generate the heel clearance pattern finally. Then we can get the complete pattern during swing phase and also finish the natural walking pattern generation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003503_s40684-021-00335-6-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003503_s40684-021-00335-6-Figure1-1.png", "caption": "Fig. 1 Leg model representation", "texts": [ "\u00a02 dynamic modeling of exoskeleton is presented. In Sect.\u00a03, the controller design is deliberated. In Sect.\u00a04, parameters for simulation study are described. Sect.\u00a05 explains the collaborative simulation (co-simulation) technique, Sect.\u00a06 presents the simulation results and discussions. Section\u00a07 provides the conclusion and the contribution of this work. The kinematics for the three joints such as hip ( 1) , knee ( 2 ) and ankle ( 3 ) of a lower extremity exoskeleton model [26, 27] is given in Fig.\u00a01. Swing phase and stance phase are the two phases of human walking cycle as shown in Fig.\u00a03. The dynamic model of the system is recognized by considering Legrangian formulation. The Legrangian L is defined by. where j = 1,2,3 the three joints, j and ?\u0307?j are the angular position and angular velocity of corresponding joints of the exoskeleton. K ( \ud835\udf03j, ?\u0307?j ) and P( j) are the total kinetic and potential energy stored in the dynamic system. From Legrangian, equation of motion of the dynamic system is given by [27], The generalized external forces acting on the system is the combination of \u0393j and vj j in the Eq", " The generalized equation for the tri-link (hip-knee-ankle) assembly of the lower extremity exoskeleton, is formed as: [27]. where = [ 1 2 3 ]T , are the angular motion of hip, knee and ankle of the exoskeleton respectively, M \u2208 \u211c3\u00d73 is the inertial properties matrix, Q \u2208 \u211c3\u00d73 is the Coriolis and centripetal matrix, N \u2208 \u211c3\u00d71 is the gravity vector and \u0393 \u2208 \u211c3\u00d71 is the torque applied to the three joints of the system. From Eq.\u00a0(3) ?\u0308? is obtained including the uncertainty,w(t)\ufffd as, Equation\u00a0(4) is expanded as, The elements of the above equation are obtained (Ref [28]) from the schematic of the exoskeleton shown in Fig.\u00a01. The dynamic model of the exoskeleton can be obtained by considering the leg in stance phase as an inverted pendulum model, and is given as where ?\u0308? = [ ?\u0308?1 ?\u0308?4 ] \u0393\ufffd 3 \u2014right ankle torque during stance phase \u0393\ufffd 6 \u2014left ankle torque during stance phase FGRF is the ground reaction force which will be considered during the stance phase of the leg. (3)M(\ud835\udf03)?\u0308? + Q ( ?\u0307?, \ud835\udf03 ) ?\u0307? + N(\ud835\udf03) = \u0393 (4)?\u0308? = M\u22121(\ud835\udf03) [ \u0393 \u2212 Q ( ?\u0307?, \ud835\udf03 ) ?\u0307? \u2212 N(\ud835\udf03) \u2212 w(t) \ufffd ] (5) \u23a1 \u23a2\u23a2\u23a2\u23a3 ?\u0308?1 ?\u0308?2 ?\u0308?3 \u23a4 \u23a5\u23a5\u23a5\u23a6 = \u23a1 \u23a2\u23a2\u23a2\u23a3 m(1,1) m(1,2) m(1,3) m(2,1) m(2,2) m(2,3) m(3,1) m(3,2) m(3,3) \u23a4 \u23a5\u23a5\u23a5\u23a6 \u22121 \u239b\u239c\u239c\u239c\u239d \u2212 \u23a1\u23a2\u23a2\u23a2\u23a3 q(1,1) q(1,2) q(1,3) q(2,1) q(2,2) q(2,3) q(3,1) q(3,2) q(3,3) \u23a4\u23a5\u23a5\u23a5\u23a6 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003660_j.asr.2021.06.018-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003660_j.asr.2021.06.018-Figure1-1.png", "caption": "Fig. 1. Dynamic components a", "texts": [ " The paper structure includes: i) the elaboration of the dynamic model and the control of the free gyroscopic rotor in magnetic suspension, as well as the nonlinear dynamic model of the DGMSCMG taking into account the interconnections of its dynamic components (AMB-rotor, inner gimbal, and outer gimbal) \u2013 Section 2; ii) the design of the automatic control systems associated to the nonlinear models of the DGMSCMG dynamic components by using the concept of dynamic inversion and neural networks \u2013 Section 3; iii) the software validation of the new control systems and the analysis of the obtained results \u2013 Section 4; iv) conclusions of the paper and ideas for future work \u2013 Section 5. The dynamic components of DGMSCMG are highlighted in Fig. 1.a, while the graph of their interdependencies is given in Fig. 2. OX iY iZi is the inertial frame, Ox0y0z0 the local orbital frame, Oxbybzb base (satellite) tied reference frame, Oxsyszs support frame tied trihedral (position \u2018\u2018zero\u201d of the outer gimbal), Oxeyeze outer gimbal tied reference frame, Oxiyizi inner gimbal tied reference frame, Ox1y1z1 gyroscopic stator tied reference frame, and Oxryrzr Ox2y2z2 gyroscopic rotor tied reference frame. Notations: All over the paper, the physical vectorial variables are represented under the geometric form as: ~X ; with ~X \u00bc X 1 i ", " 2, x0 is the orbital speed of the satellite, xb the absolute angular rate of the satellite, x0b the satellite angular rate relative to local orbital frame, xbs the angular rate of the support (s) relative to the base (b), xse \u00bc _re the angular rate of the outer gimbal (e) relative to the support, xei \u00bc _ri the angular rate of the inner gimbal (i) relative to the outer gimbal, xi1 the angular rate of the stator relative to the inner gimbal, xr the angular rate of the rotor relative to the stator, Xr the absolute angular rate of the gyroscopic rotor, Xi the absolute angular rate of the inner gimbal, Xe the absolute angular rate of the outer gimbal,M r the outer torque of the gyroscopic rotor acting on it, M i the torque directly applied to the inner gimbal, Me the torque directly applied to the outer gimbal. Because the support frame is considered fixed on the base (satellite) and the gyroscopic stator is fixed on the inner gimbal, we write: xbs \u00bc 0 and xi1 \u00bc 0 ; respectively. Fig. 1.b presents the rotations of the gyroscopic rotor relative to the stator; according to this figure, the angular rate vector of the stator is: x!r \u00bc _a !\u00fe _b ! \u00fe _c ! \u00f01\u00de The interactions between the rotor and the DGMSCMG gimbals are initially omitted, i.e. the gyroscope is nd their associated frames. Fig. 2. Graph of the interdependencies associated to the dynamic components. considered to be free. In addition to the highlighted angular displacements, the rotor can also move linearly along its axes, with xr and yr", " (18), the above-presented expression of Xr = i becomes: Xr = i \u00bc xr \u00fe xei \u00fe Ai exse \u00fe Ai bxb \u00bc xr = i \u00fe xei = i \u00fe xse = i \u00fe xb = i ; \u00f020\u00de thus Xr = i is the sum of the projections of the angular rates xr ; xei ; xse ; and xb on the inner gimbal tied frame\u2019s axes. The angular rates from Eq. (20) have the algebraic forms: xr \u00bc _a _b _c T ; xei \u00bc _ri \u00bc _ri 0 0\u00bd T ; xse \u00bc _re \u00bc 0 _re 0\u00bd T ; xb \u00bc xb xb xb yb xb zb T ; \u00f021\u00de the components from Eq. (21) are the projections on the axes of the rotor tied frame (Oxryrzr), on the axes of the inner gimbal tied frame (Oxiyizi), on the axes of the outer gimbal tied frame (Oxeyeze), and on the axes of the base tied frame (Oxbybzb), respectively. According to Fig. 1.b and Fig. 4, since a and b have very small values, it results: Xr = i \u00bc Xr xi Xr yi Xr zi 2 64 3 75 \u00bc _a\u00fe _ri \u00fe xb xi _b\u00fe _recosri \u00fe xb yi _c _resinri \u00fe xb zi 2 64 3 75 ; \u00f022\u00de xb xi \u00bc xb xbcosre xb zbsinre ; xb yi \u00bc xb xbsinricosre \u00fe xb ybcosri \u00fe xb zbsinricosre ; xb zi \u00bc xb xbcosrisinre xb ybsinri \u00fe xb zbcosricosre : \u00f023\u00de The rotor absolute kinetic torque and the one projected on the axes of inner gimbal tied frame are: K r \u00bc J rXr ; K r ; i \u00bc J rXr = i ; with J r \u00bc diag J rx J ry J rz\u00bd T ; J rx \u00bc J ry << J rz : The absolute vectorial equation of the rotor\u2019s kinetic moment is (Lungu et al", " i \u00fe J ryXr yi j ! i \u00fe J rzXr zi k ! i : By projecting eq. (24) on the axes of the Oxiyizi frame and taking into account that J rx \u00bc J ry << J rz ; one gets: Fig. 4. Rotations of Oxiyizi frame relative to Oxeyeze frame (a); rotations of Oxeyeze frame relative to Oxbybzb frame (b). J rx _Xr xi \u00fe J rzXr ziXr yi \u00bc Mr xi ; Mr xi ffi Mr xr \u00bc Mc xr ; J rx _Xr yi J rzXr ziXr xi \u00bc Mr yi ; Mr yi ffi Mr yr \u00bc Mc yr ; J rz _Xr zi \u00bc Mr zi ; Mr zi \u00bc 0 ; \u00f025\u00de one considered that a and b have small values and, according to Fig. 1.b, the frames Oxryrzr and Oxiyizi are almost overlapping. Now, replacing eq. (22) into (25), we obtain: J rx \u20aca\u00fe \u20acri \u00fe _xb xi\u00f0 \u00de \u00fe Kr zi _b\u00fe _recosri \u00fe xb yi \u00bc Mc xr ; Jrx \u20acb\u00fe \u20acrecosri _ri _resinri \u00fe _xb yi Kr zi _a\u00fe _ri \u00fe xb xi\u00f0 \u00de \u00bc Mc yr ; J rz \u20acresinri \u00fe _ri _recosri _xb zi\u00f0 \u00de \u00bc 0 ; Kr zi ffi Jrz _resinri \u00fe xb zi\u00f0 \u00de : \u00f026\u00de J ix \u20acri \u00fe _xb xi\u00f0 \u00de \u00fe J i z J i y _recosri \u00fe xb yi _resinri \u00fe xb zi\u00f0 \u00de \u00bc Mxi \u00feMc xr ; J iy \u20acrecosri _ri _resinri \u00fe xb yi \u00fe J i x J i z\u00f0 \u00de _resinri \u00fe xb zi\u00f0 \u00de _ri \u00fe xb xi\u00f0 \u00de \u00bc Myi Mc yr ; J iz \u20acresinri _ri _recosri \u00fe _xb zi\u00f0 \u00de \u00fe J i y J i x _ri \u00fe xb xi\u00f0 \u00de _recosri \u00fe xb yi \u00bc Mzi : \u00f029\u00de Concluding, the dynamic model of the AMB-rotor in outer gimbal suspension is described by eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001235_j.acme.2014.11.003-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001235_j.acme.2014.11.003-Figure4-1.png", "caption": "Fig. 4 \u2013 Pressure distribution on the gear circumference.", "texts": [ " for pumps with radial clearance compensation. The size of the pressure build-up area also depends on the operating parameters of the pump [2]. In the calculation of the radial forces the size of areas w1 and w2 can be established based on the measured pressure distribution. The changes of the position of the sealing point cause periodic changes in the pressure in the range of w3'. After the trapped volume is connected with the suction chamber, the sealing point changes its position from point S to P (Fig. 4), which causes a sudden change in the radial forces. The resulting radial forces are obtained by integrating over the circumference. The pressure forces on the driving and driven gear can be calculated from: Fhx1 \u00bc p b ra sin F2 \u00fe te 2 cos aw Fhy1 \u00bc p b ra cos F2 \u00fe rb te 2 sin aw Fhx2 \u00bc p b ra sin F2 \u00fe te 2 cos aw Fhy2 \u00bc p b ra cos F2 \u00fe rb te 2 sin aw (6) In Fig. 5 the calculated time courses of the radial forces for both gears have been shown. The courses of the components of the pressure load in the x-direction are identical for both gears; however, they have an opposite sign" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002530_978-3-030-42006-2-Figure6.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002530_978-3-030-42006-2-Figure6.2-1.png", "caption": "Fig. 6.2 Robot frames", "texts": [ " Let Fb, Fc, Fo be the coordinate frames attached to the robot base, to the camera and to the object, while T b o represents the homogeneous transformation between the \u00a9 Springer Nature Switzerland AG 2020 C. Copot et al., Image-Based and Fractional-Order Control for Mechatronic Systems, Advances in Industrial Control, https://doi.org/10.1007/978-3-030-42006-2_6 99 100 6 Simulators for Image-Based Control Architecture Fo and Fb and the homogeneous transformation between Fc and Fb is given by T b c , Fig. 6.2. The image-based control architecture presented in Fig. 6.1 has as input a set of n points defined in the Cartesian space, i.e. points which describe the object (block \u2018oP\u2019). Knowing the starting position and orientation of the object with respect to the robot base (block \u20180To\u2019), the coordinates of the points with respect to the robot base (block \u20180P\u2019) can be computed using homogeneous transformation. Starting from the hypothesis that the initial and the desired position and orientation of the camera are known, the coordinates of the initial point (cP), respectively, the coordinates of desired points (cP\u2217) can be detected with respect to camera coordinate system by using homogeneous transformations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003664_tia.2021.3084549-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003664_tia.2021.3084549-Figure6-1.png", "caption": "Fig. 6. Field magnetic fluxes path.", "texts": [ " low torque levels, but a large field weakening current causes an increase in the copper loss of the field winding and a decrease in total efficiency. Therefore, a very small field weakening current is applied only in the high-speed region where the iron loss increases. In the proposed motor, the rotor support component uses SUS304 (a nonmagnetic material), the ferrite PM uses NMF12G+ (Hitachi Metal, Ltd.), the SMC uses ML35D (Kobe Steel, Ltd.), and the stator core uses electromagnetic steel sheet 35JN230 (JFE Steel Corp.). Table II lists the material properties of NMF-12G+, ML35D, and 35JN230. Fig. 6 shows the path of the field magnetic flux in the whole motor. The shaft and the motor case lid made of carbon steel are used as the magnetic path of the field magnetic flux in order to efficiently excite the SMC. It is possible to increase the torque density by designing the shaft and the case cover of the motor so as to serve as both a structural component and the field magnetic path. Fig. 7 shows the magnetic flux density distribution in the air gap by 3D-FEA using the JMAG-Designer simulation software (JSOL corporation) under no field current, maximum field strengthening, and maximum field weakening" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003279_j.matpr.2020.12.094-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003279_j.matpr.2020.12.094-Figure7-1.png", "caption": "Fig. 7. Maximum and minimum strain values at three distinctive laser power (a) 25 W (b) 50 W (c) 75 W.", "texts": [], "surrounding_texts": [ "In this research the preliminary experimentations, RSM-PSO based optimization and thermal analysis followed by stress, strain and deformation results were studied for ES-LBTM approach during the generation of hole profile. It has been observed that for the material ablation characteristics, laser power acts as a dominant parameter out from the other parameters which is studied both from experimentations and simulation results. Moreover, from this research it has also been observed that the laser current intensity promotes laser power to impact a lot during material ablation, with the constant time period and scanning velocity. From the RSM based PSO optimization result it has been obtained the optimal parametric setting of 28 A, 2.6 kHz and 19.28% can give MRR up to 0.0004529 g/s. With this endeavor, an attempt has been made to study the material ablation characteristics in ES-LBTM approach. This research further implied to do more numbers of experimentations followed by analysis of thermal stress and its impacts as it alters depending on the thermal conductivity and thermal expansion coefficient of the material; which in turn widen up the areas in advanced manufacturing approach towards the context of ample range of industrial applications with DSS super alloy. Declaration of Competing Interest The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper." ] }, { "image_filename": "designv11_22_0001850_amr.1137.61-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001850_amr.1137.61-Figure2-1.png", "caption": "Fig. 2 Elements of ultrasonic machine (AP-1000) [4]", "texts": [ " For efficient machining to take place, the tool and horn must be designed with consideration given to mass and shape so that resonance can be achieved with in frequency range capability of the ultrasonic machine. Basic components of ultrasonic machine tool Ultrasonic machine comprises of number of essential elements which are necessary to make the working of machine proper. These major components are ultrasonic power supply, transducer or converter, ultrasonic horn, tool and tool assembly etc. Fig. 2 shows the basic element of ultrasonic machine set-up. These all the components are described below; In Ultrasonic machine the power supply is more accurately characterized as a high power sine wave generator that offers the user control over the frequency and power of the generated signal. It converts low frequency (50-60 Hz) electric power to high frequency (20-25 kHz). This electric signal is applied to transducer which further converts it into vibration. Transducers convert the electrical energy into mechanical energy" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003530_s11740-021-01050-6-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003530_s11740-021-01050-6-Figure7-1.png", "caption": "Fig. 7 Laser scanning strategy: a part\u2019s isometric view, b support material system, c laser scanning strategy top view, d fine tuning parameters top view (upskin filling and borders)", "texts": [ " The part has the following dimensions: 32.71 \u00d7 21.23 \u00d7 190\u00a0mm. Other specifications are included in Table\u00a07. In Table\u00a08, the process parameters and material specifications are listed. They have been selected from different screening experiments and taking into account the literature review provided in Table\u00a03. The laser scanning strategy selected for the fabrication of this product was quad island (meander) in which the laser is divided into small squares and fills the area in a different order. In Fig.\u00a07, the implant is shown with its support structures (Fig.\u00a07a, b), and the laser scanning strategy filling the quad islands in a specific order (filling, upskin filling, contours, borders, upskin border) (Fig.\u00a07c\u2013i). Consumed power was quantified through time using an energy logger model Fluke 1732 (Fluke Corporation, Everett, WA, USA). The data was analyzed and identified in Fluke Energy Analyze Plus Software. The Fig.\u00a08 presents the power profile and time breakdown resembling the generic schematics shown in Fig.\u00a04. The power averages were obtained from the measurements and analyzed with the previous equations to present the LCI data. L-PBF process can be divided into three main sections, in agreement with the energy breakdown of the generic model presented above" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000457_00207179.2019.1669826-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000457_00207179.2019.1669826-Figure3-1.png", "caption": "Figure 3. Schematic diagram of a magnetic levitation system.", "texts": [ " (33) According to the final value theorem, there is lim t\u2192\u221ee (t) = lim s\u21920 sE (s) = lim s\u21920 (sI \u2212 Ac) \u22121 BB+ \u00b7 s (1 \u2212 Gf (s) ) Ud (s) = 0. Therefore, the closed-loop system achieves zero steady-state tracking error. In Theorem 6.2, the conditions in Equation (29) could be satisfied while appropriate filter Gf (s) is selected. The internal model principle could be adopted for the design of Gf (s) to achieve asymptotic reference tracking and disturbance rejection (Ren et al., 2017). The schematic diagram of a MagLev system is displayed in Figure 3. The system is controlled by the voltage u, which introduces a current i in the electromagnet to build a magnetic field. The ball position q can be regulated by changing the voltage amplitude. The linearised state-space model of a MagLev system is represented as follows (Goodall, Michail, Whidborne, & Zolotas, 2009): { x\u0307 = Ax + Bu + A \u00b7 x + d y = Cx (34) where x = [ q q\u0307 i ]T is the system state, u is the control input (voltage) and y = q is the system output; A and d denote the model uncertainties and external disturbances, respectively, and letud = A \u00b7 x + d denote the general uncertainties and disturbances; A, B and C denotes the systemmatrix, input matrix and output matrix, respectively, which are represented as A = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 0 1 0 2Kf I2o MsG3 o 0 \u22122Kf Io MsGo 0 0 \u2212Rc Lc + KbNcAp/Go \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 B = [ 0 0 1 Lc + KbNcAp/Go ]T C = [ 1 0 0 ] The physicalmeanings and values of theMagLev systemparameters are given in Table 1 (Goodall et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003454_j.optlastec.2021.107058-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003454_j.optlastec.2021.107058-Figure5-1.png", "caption": "Fig. 5. Finite element model of quasi-static tensile specimen.", "texts": [ " Based on the aforementioned J-C model, the deformation behavior of Hastelloy X in the quasi-static tensile process are simulated via ABAQUS software. The finite element model of uniaxial tensile experiment was established. According to literature [23], the density, poisson\u2019s ratio and elasticity modulus of Hastelloy X are 8220 kg/m3, 0.3, and 205 GPa. Thus, substitute these material properties parameters and the constitutive model parameters listed in Table 4 into the software, the simulated result will obtain. In this model, while fixing one end of the sample, a displacement boundary condition is applied to the other end of the sample (Fig. 5). The consistency between simulation and test was achieved by setting the same loading displacement and calculation time as those of the tests. The finite element model is discretized using reduced integration (C3D8R) eight-node hexahedral elements. In order to ensure the calculation accuracy, the mesh in the gauge length of the stretched rod is refined, and the total element number is 10640. The simulated stretching process is shown in Fig. 6. In the initial stage (Fig. 6a), the specimen deforms uniformly in the gauge length section, and then the plastic deformation tends to concentrate in the middle until the specimen fractures" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure7.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure7.2-1.png", "caption": "Figure 7.2. Tipped lathe tools", "texts": [ " XX H (grey) \u2013 hard steels and chilled cast iron. Subsequently the various grades of insert are available with corresponding colour coding to provide ease of choice of application. Cemented carbides are used in cutting tools for turning, boring, milling, drilling etc. Solid carbide drills, taps and milling cutters are available for cutting a wide range of materials. Cemented carbides are also available in the form of tips brazed to a suitable tool shank, or as an insert clamped in an appropriate holder Fig. 7.2. The inserts are produced by mixing the metal powders in the correct proportions, pressing them to the required shape under high pressure and finally heating at a temperature in the region of 1400\u00b0C, a process known as sintering. The sintering stage D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 7 Cutting tools and cutting fluids 104 7 resistance and reduced smearing tendencies (tendency of the workpiece material to smear or cling to the surface of the tool)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000190_s11044-019-09681-5-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000190_s11044-019-09681-5-Figure11-1.png", "caption": "Fig. 11 Butterfly trajectory response of the end effector in task space", "texts": [ " The desired trajectory of the end effector as shown in Fig. 10 represents a butterfly path in which the X and Y coordinates are described by pd(t) = [ xd yd ] = [ 0.02 cos(t)ecos(t) \u2212 2 cos(4t) \u2212 sin5( t 12 ) + 0.35 0.02 sin(t)ecos(t) \u2212 2 cos(4t) \u2212 sin5( t 12 ) + 0.2 ] (51) The initial posture of the manipulator is assumed at q(0) = [\u03c0 \u2212\u03c0 2 \u2212\u03c0 2 ]T (rad.) that corresponds to the end-effector initial pose: p(0) = [01]T m. The uncertainties including external disturbances and parameter variations are all applied at t = 5.0 s. Figure 11 shows the trajectory response of the end effector. Figure 12 shows the tracking position error of the end effector in the X-axis and Fig. 13 shows the profile of the tracking position error of the end effector in the Y -axis. It is obvious that the ANFIS-AIDCSMC eliminates the chattering phenomenon and yields a favorably better tracking response with small steady state error. The different errors are tabulated in Table 2. The proposed algorithm ANFIS-AIDCSMC was tested for the operational space control of a rigid three-link revolute manipulator model [9] which is shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000439_j.mechmachtheory.2019.103612-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000439_j.mechmachtheory.2019.103612-Figure7-1.png", "caption": "Fig. 7. Forces diagrams of the (a) arm and (b) forearm.", "texts": [ " 6 , M 2 and M 3 are the net joint moments at the shoulder and elbow, respectively; F 2 is the vertical force at shoulder, which corresponds to the force usually measured with force plates located under the user\u2019s back [21,55] ; and F T is the vertical force acting on the user\u2019s wrist. In the model the net joint moment at the wrist has been neglected. It must be noted that the mechanism is overactuated by the moments at the shoulder and elbow. This mean that there is an infinite number of combinations of M 2 and M 3 that allows lifting the load along the same trajectory. Fig. 7 shows the free body diagrams of the arm and the forearm, where the mass and rotational inertia of the bodies and the internal and external forces at joints have been included. According to Fig. 7 (a), the equilibrium equations for horizontal forces, vertical forces and moments with respect to the CG of the arm read as: F 12 ,x \u2212 F 32 ,x = m a \u0308r ga,x (15) F 2 \u2212 F 32 ,y = m a \u0308r ga,y + m a g (16) F 12 ,x L ap sin \u03b8a + F 32 ,x L ad sin \u03b8a \u2212 F 2 L ap cos \u03b8a \u2212 F 32 ,y L ad cos \u03b8a + M 2 \u2212 M 3 = I a \u03b8\u0308a , (17) where L ap and L ad are the proximal and distal distances from the center of gravity of the arm, respectively, I a is the moment of inertia of the arm with respect to its CG and g is the gravity acceleration. Similarly, from Fig. 7 (b), the equilibrium equations for the forearm are written as: F 23 ,x \u2212 F 43 ,x = m f r\u0308 g f,x (18) F 23 ,y \u2212 F T = m f r\u0308 g f,y + m f g (19) F 23 ,x L f p sin \u03b8b + F 43 ,x L f d sin \u03b8b \u2212 F 23 ,y L f p cos \u03b8b \u2212 F T L f d cos \u03b8b + M 3 = I f \u03b8\u0308b , (20) where L fp and L fd are the proximal and distal distances from the center of gravity of the forearm, respectively, and I f is the moment of inertia of the forearm with respect to its CG. The term of force F T includes the force due to the resistance system F R , the force due to the hand mass m h and the friction F f that appears at the vertical guides of the Smith machine [55] as follows: F T = F R + m h (g + y\u0308 ) + F f (21) Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001283_cse.2014.104-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001283_cse.2014.104-Figure8-1.png", "caption": "Fig. 8. AR.Drone.", "texts": [ " If t is the control time, y(t) is the control output, and e(t) is error between reference input and control output, the PID control then can be expressed as follow: (4) where Kp is proportional gain, Ki is integral gain, and Kd is derivative gain. In the proposed teaching material, students are required to properly set these parameters in order to obtain a desired response of the AR.Drone. Since the yaw angle and the altitude of the AR.Drone need to be controlled in order to track the target color, six parameters are required to be set during the learning process. IV. EXPERIMENT A. AR.Drone The AR.Drone used in the proposed teaching material is shown in Fig. 8. It is a product of Parrot Company, and it costs around 300 USD. It is equipped with two cameras (one at the front and another at the bottom), an ultrasonic sensor, and an accelerometer. The cameras can allow user to screen the views, and the ultrasonic sensor and the accelerometer can provide user altitude, yaw, pitch, and roll during flying. Basically, the AR.Drone only provides piloting capabilities via iOS devices. However, by setting ad hoc connection, and using the software library (AR.Drone SDK), it is possible to pilot it via a computer" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure6-1.png", "caption": "Fig. 6. Evolution of a test rig design for nodes in gridshell structures.", "texts": [ " For numerical simulation purpose, the average of the test data is employed and simplified by a multilinear curve. Fig. 8 shows the averaged result from the tensile tests, and the simplified stress-strain curve used for non-linear simulations. The printing direction of dog-bone sample is also shown in Fig. 5. The design of test rig is evolved during a process in which the main idea of the final design is generated based on the necessities that the final design is capable of testing nodes under various loading conditions, such as bending, shear, axial and combined loads. Fig. 6 shows three test rigs generated during the design process. The first test rig is proposed for a symmetrical three-way node designed for resisting outof-plane bending moment. As shown in Fig. 6(a), the upper and lower parts of the test rig are designed to be clamped in the hydraulic jack. In this test rig, the out-of-plane bending moment is generated on each connecting member by applying two vertical forces in the opposite directions and with a certain eccentricity. This method may not be appropriate for full scale node test due to the limited space inside the testing machine. Besides, the test rig is customised for only one node under one loading case. In this design, the upper and lower parts of the test rig are complicated to be manufactured. In the following iterations shown in Fig. 6(b) and (c), the testing system is fixed onto the ground in order to eliminate the dependency on the size of testing machine. Also, to produce different bending moments, multiple bolt holes are considered in the connecting plate so that vertical loads can be applied in different locations. In the final design of the test rig (Fig. 6(d)), the distances between bolt holes and the dimensions of the connecting plates are determined in a modular way in which different loading conditions are obtained by changing the configuration of plates and bolts. Configurations of the test rig for six different loading conditions, including bending, shear, compression, tension, and two combined loading conditions are shown in Fig. 7. To pre-determine the load resisting capacities of the designed structural nodes, non-linear finite element analysis is carried out using Abaqus" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002379_0954406220917424-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002379_0954406220917424-Figure3-1.png", "caption": "Figure 3. Dynamic model of the planetary system.", "texts": [ " When considering the axial and profile modifications, the tooth surface equation15 is x0 \u00bc r0b sin uk \u00fe 0 00\u00f0 \u00de uk cos uk \u00fe 0 00\u00f0 \u00de\u00bd y0 \u00bc r0b cos uk \u00fe 0 00\u00f0 \u00de \u00fe uk sin uk \u00fe 0 00\u00f0 \u00de\u00bd y0 Ra \u00fe s0a=2 2 \u00fe z0 B=2\u00f0 \u00de 2 \u00bc R2 a 8>< >: \u00f07\u00de Thus, the component of the offset along the meshing line caused by the micro-modification is h \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x2 \u00fe y2 p sin k cos \u00fe z sin ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x02\u00fe y02 p sin k cos \u00fe z0 sin h i \u00f08\u00de where k is the pressure angle on the involute. Dynamic modeling and analysis of helical planetary gear set The dynamic model of this system shown in Figure 3 employs a number of simplifying assumptions: 1. Each gear body is assumed to be rigid while elastic deformations of the planetary carrier and gears are ignored. 2. Each gear mesh interface is simplified with a spring having a periodically time-varying stiffness and viscous damping elements along the gear LOA. 3. The relative sliding and sliding friction in the meshing teeth are not considered. 4. The mass, moment of inertia, radius, and mean meshing stiffness of each planet gear are identical. In Figure 3, xi, yi, zi and uix, uiy, uiz (i\u00bc s, c, r, n) are the elastic deformations and torsional deformations around the axes of each component, respectively. kix, kiy and cix, ciy are the bearing stiffness and damping of the corresponding component in the x, y directions, respectively. kju and cju (j\u00bc s, c, r) represent the torsional stiffness and damping of s, c, and r, respectively. krn, ksn and crn, csn denote the mesh stiffness and damping of the nth sun\u2013planet (s\u2013p) and planet\u2013ring (p\u2013r) gear pairs, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003796_s10846-021-01450-x-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003796_s10846-021-01450-x-Figure3-1.png", "caption": "Fig. 3 Schematic diagram of how to realize fully-actuated movement for the novel multirotor UAV", "texts": [ " (16) From the assumption A4, it is obvious that the matrix A4A T 4 will be a positive definite matrix, therefore, A4 will be a row full rank matrix. It\u2019s because of this property that the novel multirotor can move independently with 6-DOF, therefore, we can develop fully actuated controller by using um as control input. It should be noted that in order to realize this purpose, we need to use at least eight rotors, for example, if we use a similar 6-rotor structure, the corresponding matrix A4 will not be a row full rank matrix, therefore it will be difficult to realize fully actuated control with similar structure. In Fig. 3, the way of how to realize 6-DOF independent movement for the novel multirotor UAV is shown. In Fig. 3a, b and c, they show the change of the rotor speeds in the cases that this novel multirotor moves along the positive direction of the body coordinate OB , respectively, where \u201c+\u201d means that speed increases, and \u201c-\u201d means that speed decreases. In Fig. 3d, e and f, they show the change of the rotor speeds in the cases that this novel multirotor rotates around the positive direction of the body coordinate OB , respectively. It is obvious that the opposite speed change will make the multirotor moves along or rotates around opposite direction of the body coordinate OB , respectively. Therefore, the novel multirotor can realize fully-actuated movement. Although we use the structure design of 8 rotors, different from the proposed structure design in [25\u201331], this novel structure is symmetrical, which is conducive to control design and operation, and does not need servo structure, which reduces the complexity of the structure" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000305_012001-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000305_012001-Figure1-1.png", "caption": "Figure 1. Basic FDM process [11]", "texts": [ " Keyword: FDM, PLA (Polylactic acid), PLA-Copper, Compressive Strength, ANOVA, RSM, FCCCD. The FDM technology, one of the most commonly used AM methods nowadays due to its simplicity in structure, wide variations in equipment design that suitable for all types of users, from the small design machine for normal using for heavy duty one for fabricating prototype for industrial using. It works on a layer-by-layer principle, the material is liquefied and then laid down in layers on top of each layer until the part is formed (Fig 1). The material used in this case is in a form of a wire filament, there's a wide variety of filament type range from pure plastic like PLA or ABS to plastic with a solid powder composite materials like PLA-copper, ABS-wood, etc. Along with its development, many problems arise, such as: how to optimize fabricated the technical requirements of FDM products; reduce fabricating times and improve quality of the finished parts, and etc. Many solutions are offered, likes developing a new kind of materials or correcting process parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001142_sii.2014.7028137-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001142_sii.2014.7028137-Figure6-1.png", "caption": "Fig. 6. DD motor and ER brake", "texts": [ " It is thus not suitable for precise speed control. However, we believe that there is potential in controlling the speed of the motor by the resistance force of the fluid because of the fast response of the ER fluid. In this study, we experimentally determined the relationship between the torque T and electric field E of the ER fluid. The relationship is shown in Fig. 7. Fig. 7 shows the torque transmitted by the ER fluid for an electric field of 0\u20132 kV. The torque was measured by a sensor as arranged in Fig. 6. Equation (1) is a quadratic approximation of the relationship. T = 0.0364E2 + 0.0901E + 0.0066 (1) Preliminary experiments were carried out to determine the response to different motor speeds. Here, an I-P control plant was used to obtain (1) [8]. Fig. 8 is a block diagram of the I-P speed control. The target angular speed was 2.0 rad/s and control began 1 s after the motor started rotating. ERB in Fig. 8 refers to the control box of the ER fluid brake. The relationship between torque \u03c4 and gains KI and KP is described by (2) for I-P control of speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000342_j.triboint.2019.105881-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000342_j.triboint.2019.105881-Figure4-1.png", "caption": "Fig. 4. Oil film section.", "texts": [ " The calculation mesh zones were divided in the three parts: the oil film, the oil and the pad. The influence of the grid size on the calculation results was very small in the X direction. In the Y direction, especially for the oil film, selecting the number of grid divisions was very important, as shown in Table 2 for Fig. 3(a). In total, five divisions of the oil film in the Y direction were used for the condition of guaranteeing accuracy [21]. TEHD analysis involves Reynolds equation and energy equation. According to 3D-TEHD theory, for the oil film in Fig. 4, the gov- erning equation of its pressure is Reynolds equation: \u239c \u239f \u2202 \u2202 \u239b \u239d \u22c5 \u2202 \u2202 \u239e \u23a0 + \u2202 \u2202 \u239b \u239d \u22c5 \u2202 \u2202 \u239e \u23a0 = \u22c5 \u2202 \u2202x F p x y F p y U z x m 2 2 (7) Where: \u222b= \u22c5 \u2212F dzh x z z z z \u03b72 0 ( , ) ( )m , \u222b \u222b=z dz dz/m h x y z \u03b7 h x y \u03b70 ( , ) 0 ( , ) 1 The pressure boundary condition is: The boundary pressure around the pad is 0. The governing equation of oil film temperature is: \u239c \u239f\u22c5 \u22c5\u239b \u239d \u22c5\u2202 \u2202 + \u22c5\u2202 \u2202 + \u22c5\u2202 \u2202 \u239e \u23a0 \u2212 \u22c5\u2202 \u2202 = \u23a1 \u23a3\u23a2 \u239b \u239d \u2202 \u2202 \u239e \u23a0 + \u239b \u239d \u2202 \u2202 \u239e \u23a0 \u23a4 \u23a6\u23a5 \u03c1 u T x v T y w T z \u03bb T z \u03b7 u z v z C 2 2 2 2 (8) The temperature boundary condition is: = \u2265T y z T u y z(0, , ) [ (0, , ) 0] (9-1) =T x y T( , , 0) r (9-2) \u22c5\u2202 \u2202 = \u22c5 \u2202 \u2202= = \u03bb T z \u03bb T zz h p p z 0p (9-3) \u2212 \u239b \u239d \u2202 \u2202 \u239e \u23a0 = \u22c5 \u2212\u03bb T n h T T( )p \u03b1 c \u03b1 s (9-4) Formula (9-1) represents the temperature condition of the oil flow in the loading area" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002952_icee50131.2020.9261035-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002952_icee50131.2020.9261035-Figure4-1.png", "caption": "Fig. 4. Rotary inverted Pendulum", "texts": [], "surrounding_texts": [ "Figure. 5 presents the rotary pendulum schematic diagram: Where \u03b8 is the angle of arm and\u03b1 angle of the pendulum deviating from the upright position with radian unit. , , ,\u03b8 \u03b8 \u03b1 \u03b1 are the arm and pendulum angular velocity and angular acceleration respectively. The physical definition and unit of parameters are listed in Table 1. TABLE I. THE PHYSICAL DEFINITION OF PARAMETERS OF THE SYSTEM Parameter Definition Parameter Definition m Integrate mass of the pendulum g Gravity coefficient M Integrate mass of the arm T Control input torque 1m Mass of the arm 1I Moment of Inertia of the arm 2m Mass of the pendulum 2I Moment of Inertia of the pendulum 1l Length of arm B\u03b1 Viscous damping constant of pendulum Authorized licensed use limited to: Tsinghua University. Downloaded on December 19,2020 at 11:32:14 UTC from IEEE Xplore. Restrictions apply. 2l Length of pendulum B\u03b8 Viscous damping constant of arm cm1l Distance of arm\u2019s centre of mass fT \u03b8 Coulomb friction constant of arm cm2l Distance of pendulum\u2019s centre of mass \u03b5 Elasticity coefficient of the arm and the input torque \u03b8 Angle of the arm \u03b1 Angle of the pendulum The Lagrangian method is used to extract the system mathematical model. The Lagrangian equation is given by: i i i i d T T V Q dt q q q \u2202 \u2202 \u2202\u2212 \u2202 \u2202 \u2202 + = (1) Where ,V T and iQ are potential energy, kinetic energy and general torques of the system, respectively. The standard representation of the system mathematical model is shown as below: M N K U \u03b1 \u03b8 \u03b1 \u03b8 + + = (2) Where M, N m K and U are: ( ) 1 2 2 2 2 2 2 2 2 1 1 2 1 2 2 2 1 2 2 1 2 2 sin cos cos cm cm cm cm cm I m l m l I m l m l l m l l I M m l \u03b1 \u03b1 \u03b1 + + + = + + ( ) ( ) 2 2 2 2 2 2 2 1 2 2 2 2 sin cos cos sin cos cm cm cm B I m l m l l N I m l B \u03b8 \u03b1 \u03b1 \u03b1\u03b1 \u03b1\u03b1 \u03b1 \u03b1 \u03b8 + + = \u2212 + 2 \u02d9 2 ( ) sin f cm T sign K m gl \u03b8\u03b5\u03b8 \u03b8 \u03b1 + = \u2212 0 T U = Since the input of the system is in voltage form, a relation between torque and the voltage can be written as below: iT av b\u03b8= + (3) Model equations can be simplified as below: 2sin cos cos A B C M C B \u03b1 \u03b1 \u03b1 += 2 sin cos cos Bsin cos F B C N E \u03b1 \u03b1\u03b1 \u03b1\u03b1 \u03b1 \u03b1 \u03b8 + = \u2212 (4) \u02d9 ( ) sin H GsignK D \u03b8 \u03b8 \u03b1 + = \u2212 0 u U I = Where 1 2 2 1 1 2 1 cmA I m l m l= + + , 2 2 2 2 cmB I m l= + , 22 1 cmC m l l= , 22 cmD m gl= , , , , , E B F B b G B H I a\u03b1 \u03b8 \u03b8 \u03b5= = \u2212 = = = and u is the input signal. If we define: 1 2 3 4 , , , x x x x\u03b8 \u03b8 \u03b1 \u03b1= = = = (5) The systems equations can be written in nonlinear state space form as follows: 21x x= ( ) ( ) ( ) 3 3 3 4 3 2 2 2 2 3 3 2 2 3 2 4 3 4 1 2 2 2 2 3 3 sin2 cos cos sin2 2 2 cos sin2 cos cos CB CD x x EC x x x x B A Bsin x C x Fx B x x x C x x Hx Gsign x Iu B B A Bsin x C x \u2212 + \u2212 = + + \u2212 \u2212 \u2212 \u2212 + + \u2212 + \u2212 43x x= ( ) ( ) ( ) 3 3 3 4 3 2 2 2 3 3 3 2 2 3 2 4 3 4 1 2 2 2 2 3 3 2 3 2 4 4 3 sin2 cos cos sin2 2 2 cos cos sin2 cos cos 1 sin sin 2 CB CD x x EC x x x B A Bsin x C xC x B Fx B x x x C x x Hx Gsign x Iu B B A Bsin x C x E D x x x x B B x \u2212 + \u2212 + + \u2212 =\u2212 + \u2212 \u2212 \u2212 + + \u2212 + \u2212 \u2212 + " ] }, { "image_filename": "designv11_22_0003657_aero50100.2021.9438399-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003657_aero50100.2021.9438399-Figure3-1.png", "caption": "Figure 3: (a) Assembly of six SMT for dPAMT, (b) HOTDOCK interfaces, before and after coupling.", "texts": [ " The Kagan angle is given by \u03be0(J1, J2) = 2 cos\u22121 max (|w|, |x|, |y|, |z|) (2) where |w|, |x|, |y|, |z| are the components of the unit quaternion qR = (w, x, y, z) for the rotation of the principal axes of the inertia tensor J2 with respect to J1. The type of connection between the mirror tiles also constraints the assembly process. Within the project PULSAR, we are using standard interfaces that guarantee mechanical, data, power and thermal connectivity between components. The interfaces selected for the project are called HOTDOCKs [19], and they are used to fix the SMT in the assembly and to manipulate the SMTs with the RAS during the assembly process. Figure 3a represents the base scenario used for PULSAR, with six mirror tiles, and multiple HOTDOCK connectors between them. Figure 3b shows the geometrical model used to represent a HOTDOCK interface as a small cylinder with four protrusions (petals) on its upper side, which guide the connection between two interfaces, thus reducing the influence of positional and rotational uncertainties during the latching process. There are some other important considerations to take into account about the geometry of the HOTDOCK interfaces. First, they present a 90\u00b0 symmetry with respect to the perpendicular axis of coupling. Since each HOTDOCK presents the four petals on the upper face, there are four possible 3 Authorized licensed use limited to: California State University Fresno" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure11-1.png", "caption": "Fig. 11. Contact stress of internal gear pair with single contact point (unit: MPa).", "texts": [ " And constraints were imposed on the degrees of freedom of all nodes on the inner surface of the pinion along the directions of UX and UZ, i.e., radial and axial directions (UX = UZ = 0). The finite element model of internal gears with single tooth is shown in Fig. 10. We analyzed the contact conditions of tooth profiles based on the setting conditions. The input torque applied to the inner surface of pinion was 200 Nm. Loading time was 1s and the time sub steps were set to 5. Stress analysis results of the internal gears with single contact point are displayed in Figs. 11-13. From the analysis results in Fig. 11, the maximum contact stress between the internal gear pair is 1429.4 MPa. The maximum stress occurs at contact point, which is located on the middle of the tooth profile. It has regular elliptical distribution along the direction of tooth width, and the distribution area has the trend of expanding to the tooth root direction. With the increase of the contact area, the contact stress will gradually decrease. The maximum von Mises stress and shear stress of the pinion with single contact point in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003692_j.nahs.2021.101060-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003692_j.nahs.2021.101060-Figure7-1.png", "caption": "Fig. 7. H\u221e performance index \u03b6\u0304min for different \u03c1 and \u03b7.", "texts": [ " (2) With the higher quantization density \u03c1, the larger upper bound of SPP \u03b5\u0304max can be obtained. In fact, when a higher quantization density is selected, the precision of the quantizer also increases, which means that the quantized signal is close to the original signal, as a result, the requirement of system parameters is reduced, so a higher upper bound of SPP is obtained accordingly. Case 2. In this section, we will study the relationships between quantization density \u03c1, sampling instant decay rate \u03b7 and H performance index \u03b6\u0304 under zero initial condition. From Table 2 and Fig. 7 (The histogram of Table 2), we \u221e min w c R can get the following results: (1) With the increase of \u03b7, the smaller \u03b6\u0304min can be obtained. It means that the bigger of sampling instant decay rate, the slower the Lyapunov function decays, and the better performance of the system can be obtained in a certain range. (2) With the bigger \u03c1, the smaller \u03b6\u0304min can be obtained. In fact, when a higher quantization density is selected, the precision of the quantizer also increases, which means that the quantized signal is close to the original signal, as a result, the requirement of system parameters is reduced, so a better performance of the system can be obtained accordingly" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002595_s13360-020-00593-4-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002595_s13360-020-00593-4-Figure1-1.png", "caption": "Fig. 1 Projectile motion under the co-planar forces of gravity, drag and lift (Magnus)", "texts": [ " We initially provide exact solution for the limiting cases of negligible Magnus force (for small rotational and translational speeds) and of negligible drag force (for relatively large velocity and spin). Then, a perturbation solution is undertaken valid for the simultaneous but weak spin and drag forces. Finally, a full solution is also supplied under the assumption of the presence of a uniform Magnus effect during the travel time of the object. Excellent agreements between the solutions from the presented formulas and from the full numerical simulations are observed. Let us consider the motion of a projectile of uniform mass m thrown into air under the gravitational force mg as depicted in Fig. 1. The projectile flying along the path with a velocity V (t), in addition to the gravitational force, is opposed with two further fundamental forces, namely the quadratic drag force Fs due to air resistance and the quadratic Magnus force Fm owing to the object\u2019s spinning lift given, respectively, by ([9\u201311,23,25] and [27]) Fs = mgk0V 2, Fm = mgk1V 2, (1) where k0 and k1 are proportionality factors of the drag and Magnus forces, the most influential parameters of the flight trajectory of the object. It is now well known that the lift or Magnus force occurs in the perpendicular direction to the axis of rotation and path of the projectile (up or down, relying upon counterclockwise (under-spin)/clockwise rotation (top-spin)), whereas the drag force is tangential to the flight path acting in the opposite direction to the motion; refer to Fig. 1 for configuration. Despite the fact that they overall depend on the translational and rotational speeds, the spin rate, the material shape (diameter, for instance), the surface roughness of the object and the fluid properties (like density) of the medium, in the present analysis we take into account an ideal situation rendering these factors being constants. Such an argument is plausible for small size objects operating under moderate spins and moderately low velocities. It is also remarked that k0 receives only positive values, while k1 may receive negative values for top-spin and positive values for under-spin action of the projectile. We reiterate that the spin is towards the motion (anticlockwise/clockwise), as seen in Fig. 1. Side spins are not permitted to deflect the motion out of the plane. Together with these assumptions, keeping in mind that the problem remains two-dimensional taking place in a fixed plane, the motion equations are due to the Newton\u2019s second law and angular momentum conservation law ([23,25] and [27]) m dV dt = \u2212mg sin \u03b8 \u2212 Fs, mV d\u03b8 dt = \u2212mg cos \u03b8 \u2212 Fm, (2) and they are accompanied with the initial speed and angle conditions V = V0, \u03b8 = \u03b80 at t = 0. (3) By means of the dimensionless velocity v(\u03b8) = V V0 , (4) it is possible to reduce the system (2)\u2013(3) to the following non-dimensional initial value problem dv d\u03b8 = sin \u03b8 + p0v 2 cos \u03b8 + p1v2 v, v = 1 at \u03b8 = \u03b80, (5) where p0 = k0V 2 0 \u2208 R+ \u22c3{0} and p1 = k1V 2 0 \u2208 R" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000433_acc.2019.8815065-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000433_acc.2019.8815065-Figure2-1.png", "caption": "Fig. 2. PAP is working so that the rings lie on a straight line of perpendicular direction. The PAP also works for aligning circular peg and hole. For more information, see the supplementary video file.", "texts": [ " In Section III, a novel mating algorithm using PAP is given, and in Section IV, the proposed PAP-based insertion algorithm is evaluated in repeatable mating and insertion experiments. Then, the validity is discussed on the basis of the experimental results. In Section V, the generality of the PAP is discussed from theoretical and experimental aspects. In Section VI, we summarize this paper and mention our future work. A. Mechanism The mechanism of the PAP is simple in spite of its lots of advantages. Fig. 2 shows that rings and shafts are aligned in line by the effect of the PAP. Two parallel shafts hang four rings respectively. The one shaft is fixed on a table and another shaft is fixed on a robot arm. The robot arm forces two shafts to come closer. When the distance between the two shafts is not close enough, the centers of rings have some displacements in tangential direction because of gravitational force. When the shafts come close enough, the displacement of each ring comes to zero, and then the center point of the rings, the center of shafts, and their contact points lie in a straight line" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003411_lra.2021.3062296-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003411_lra.2021.3062296-Figure2-1.png", "caption": "Fig. 2. Given an end-effector that provides two frictional point contactsP1P2, center-of-mass location (perpendicular the surface) and a ground contactC (left) our problem consists on finding the contact locations that are not necessarily antipodal but can result in a feasible grasp and a trajectory p(t) that leads to this feasible state. The distance between the contacts rP1P2 is fixed throughout the motion, since the end-effector has no mobility, therefore no internal forces can be applied to the object.", "texts": [ " The concept is similar to snatching [36], however since the contacts can constrain the object of certain motions, less free space and DOFs from the manipulator may be necessary to attain a successful grasp. In addition, cylindrical type objects can also be successfully grasped if their off-plane motions are restricted by the end-effector design, which can prove more difficult with the non-prehensile snatching approach. Our problem extends the earlier zero DOF end-effector work [1] regarding its initial configuration. Given a target object placed on a flat surface with gravity normal to it (Fig. 2), and an end-effector that provides two frictional point contacts P1, P2, at a fixed (non-controllable) distance rP 1P2 at opposing edges, such that the contacts can tilt the object and accelerate it away from the contact surface, the problem can be summarized by finding a trajectory p(t) in the world frame [X,Y ] such that the object can be accelerated to a final horizontal configuration while satisfying dynamic constraints that keep the object from moving relative to the contacts (attainable according to screw theory [37] if the resultant contact wrenches and the wrench of gravity are co-planar)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001638_etep.2165-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001638_etep.2165-Figure14-1.png", "caption": "Figure 14. The simulated 2P/24 SynRM in Maxwell environment.", "texts": [ " Regarding the Figures 12 and 13, computed maximum UMP is centralized on eccentricity location (around \u03b8s= 90) under SE and is rotational in dynamic case. So, eccentricity fault invigoration occurs. (D) Model validation for healthy and faulty machine modeling In order to validate the proposed model and method, the 2P/24 SynRM (case 3) and the 4P/60 SynRM (case 1) have been simulated with same properties as listed in Table I, under the same load torque in healthy condition by FEM via Maxwell.V.16 software (Ansys, Inc., Canonsburg, PA, USA) as shown in Figure 14. The Steel_1008 is considered for both rotor and stator core material in Maxwell environment as the only difference between the FEM and proposed approach. Copyright \u00a9 2015 John Wiley & Sons, Ltd. Int. Trans. Electr. Energ. Syst. (2015) DOI: 10.1002/etep The FEM is based on Maxwell\u2019s equations, which can be found in [4,5]. The relative formulas in FEM analysis can be written as illustrated in [4] by Faiz, et al. Copyright \u00a9 2015 John Wiley & Sons, Ltd. Int. Trans. Electr. Energ. Syst. (2015) DOI: 10" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure21.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure21.1-1.png", "caption": "Figure 21.1 Engine oil circulation diagram", "texts": [ " Technical drawings can vary from thumbnail sketches to illustrate a particular piece of information through pictorial drawings in isometric or oblique projection to major detail and assembly drawings. These are covered later in this chapter. Diagrams are used to explain rather than represent actual appearances. For example, an electrical circuit diagram shows the relationship of all parts and connections in a circuit represented by lines and labelled blocks without indicating the appearance of each part. Figure 21.1 is a diagram to explain the route of oil circulation in a car engine but does not go into detail of the engine itself. Exploded views are used where it is necessary to show the arrangement of an assembly in three The purpose of an operation sheet is to set out the most economic sequence of operations required to produce a finished object or process from the raw material. Although the main purpose of operation sheets is to set out the sequence of operations, they also serve a number of other very important functions: XX They determine the size and amount of material required" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003278_1350650121991741-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003278_1350650121991741-Figure1-1.png", "caption": "Figure 1. Dynamic model of the planetary gear system.", "texts": [ " Therefore, in view of this problem, on the basis of considering nonlinear factors such as time-varying meshing stiffness, backlash, and comprehensive error, a torsional dynamic model of the planetary gear system considering friction is established and the Runge\u2013Kutta numerical method is used to solve it. The bifurcation and chaos characteristics of the system with various nonlinear parameters (excitation frequency, damping, comprehensive error, load, and backlash) are analyzed through the bifurcation diagram and the largest Lyapunov exponent (LLE) diagram, and the poincar\u00e9 map, phase diagram, time history curve and fast Fourier transform (FFT) spectrum diagram are combined to further illustrate the dynamic characteristics of the system under the influence of friction. As shown in Figure 1, the torsional dynamics model of the planetary gear is established by the concentrated mass method. In this model, the meshing gears are involute spur gears, and each meshing tooth surface is connected by a spring\u2013damping system. The mass and moment of inertia of each planetary gear are the same, and the parameters such as gear backlash, comprehensive meshing error, and damping ratio of the same type of components are also the same. The three coordinate systems constructed are as follows: (1) fixed coordinate system OXY; (2) the motion coordinate system Oxy that rotates with the planet carrier, where the x-axis on the planet carrier passes through the center of the first planetary gear; (3) motion coordinate system Onxnyn (n = 1, 2, 3, " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002872_3419249.3420075-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002872_3419249.3420075-Figure3-1.png", "caption": "Figure 3: Iterations of the PCB design from the first attempt (left) to the final design (right)", "texts": [ " The technical setup of an applied conductor and LED module is illustrated in Figure 2. To seamlessly integrate SMD components into our system, we iteratively developed a flexible PCB module with two contacts. In future iterations, the results are to be scaled up to SMD packages with a higher pin count. Starting from a PCB design traditionally used in DIY interactive textiles\u2014FR4 material with enlarged tabs for the contacts with large open vias to facilitate sewing the component to the textile\u2014we used Polyimide PCB material to achieve flexibility (Figure 3.1). Using this first PCB design, we then evaluated a number of connection methods. Sewing the PCBs to the textile traces with conductive yarn and using micro rivets required fine motor skills and/or specialized tools. Soldering the PCBs directly to the conductive traces resulted in brittle connections prone to breaking. We achieved suitable connections by melting small pellets of conductive plastics (made for DIY 3D printing) between the traces and PCBs using an iron. Further iterations frayed the edges of the PCB (Figures 3", " The rougher edge structure prevents the elements from shifting between the textile layers and eliminated the need to separately contact the PCB to the traces. In the end, we achieved good electrical contact through only the edge shape and the heat-pressed adhesive on the traces. Further testing revealed that the rigid solder joints between the SMD part and PCB require additional support for long-term use in our flexible system. We increased the surface area of the Polyimide material around the solder pads to better moderate the transition from a flexible textile to a rigid solder joint (Figure 3.4). Increasing the number of holes through which the adhesive can bond through the layers and integrating the frayed edges from the previous iteration resulted in the final design of the PCB modules (Figure 3.5). In several iterations, we developed a manufacturing process that reduces the use of time, material, tools, software, and special machines as much as possible. Due to the prefabricated conductor traces, the particularly time-consuming work steps of cutting the traces and manual insulation are no longer necessary. In addition, the color-coding provides the user with assistance in contacting the PCBs. Particularly for less experienced users, this is a good orientation and helps to avoid incorrect connections" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002328_j.mechmachtheory.2020.103873-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002328_j.mechmachtheory.2020.103873-Figure14-1.png", "caption": "Fig. 14. Mises stress on wheel flank.", "texts": [ " The bending stress on pinion and wheel flanks are depicted in Fig. 11 (a) and (b). The positive maximum value is in the root of gear flank. The negative maximum value is in the working part of gear flank. Those stress values represent the tension-compression stress in meshing process. The root stress results of pinion and wheel are shown in Figs. 12 and 13 . The blue line is enveloped from black line in Figs. 12 ( a ) and 13 ( a ). Case 2: Face-hobbed hypoid gear The Mises stress for wheel in two meshing moments are demonstrated in Fig. 14 . The 2D and 3D diagrams of contact stress are shown in Fig. 15 . The bending stress on pinion and wheel flank are depicted in Fig. 16 ( a ) and ( b ). The root stress results of pinion and wheel are shown in Figs. 17 and 18 . The contact stress and bending stress of hypoid gear flank also can be checked by ISO and AGMA standards. For the face-milled hypoid gear, the ANSI/AGMA 2005-D03 and old Gleason standards were selected to check the strength. For the face-hobbed hypoid gear, the ISO/TR 13989-2 standard was used to calculate the equivalent parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.26-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.26-1.png", "caption": "FIGURE 2.26 Camber-induced side force.", "texts": [ "41) can be expressed as follows Mzr 5Dr cos\u00f0arctan\u00f0Br \u03b1r\u00de\u00de \u00f02:57\u00de with \u03b1r 5\u03b11 SHy 1 SVy Ky\u03b1 \u00f02:58\u00de Br 5QBz9 1QBz10 By Cy \u00f02:59\u00de Dr 5Fz R0 fQDz6 1QDz7 dfz 1 \u00f0QDz8 1QDz9 dfz\u00de:\u03b3 1 \u00f0QDz10 1QDz11 dfz\u00de \u03b3 j\u03b3jg cos \u03b1 sign\u00f0VCx\u00de \u00f02:60\u00de for empirical factors Qjzi, j5B, D, and i5 1, 2, 3, . . . Thus far, we have not discussed the effect of camber on the lateral tire characteristics. With only a camber angle and no slip angle, the tire tries to follow an almost circular track that is determined by the local shape of the tire cross section. The direction of motion of the wheel is forced by the vehicle velocity vector. For example, the wheel may be moving forward in a straight path. As a result, local shear stresses arise in the contact area and build up a camber force (Figure 2.26). For a motorcycle, the camber force is the major force between tire and road that prevents the tire from sliding. In the linear range, the side force can be expressed in terms of slip angle and camber angle, as follows Fy\u00f0\u03b1\u00de5C\u03b1 \u03b11C\u03b3 \u03b3 \u00f02:61\u00de with cornering stiffness C\u03b1 and camber stiffness C\u03b3, defined as C\u03b1 5 @Fy @\u03b1 \u00f0\u03b15 0; \u03b35 0\u00de \u00f02:62\u00de Cy 5 @Fy @\u03b3 \u00f0\u03b15 0; \u03b35 0\u00de \u00f02:63\u00de The ratio of camber stiffness and tire load is referred to as the normalized camber stiffness (also denoted as the camber thrust coefficient)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001620_s12206-015-0908-1-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001620_s12206-015-0908-1-Figure4-1.png", "caption": "Fig. 4. The 1/49 turbine disc segment without assembly holes.", "texts": [ " The detailed information is given below: (1) When analyzing the assembly holes, a simplified threedimensional model of the 1/6 turbine disc segment was created, and the fir-tree mortises structure were removed from the rim of the turbine disc (see Fig. 3). (2) When analyzing the fir-tree rim region, a simplified version of the three-dimensional model of the 1/49 turbine disc segment was created. The assembly hole was simplified, and the contact analysis [10] between the turbine disc and the blades was done as shown in Fig. 4. To ensure the mesh quality of the three-dimensional model and avoid excessive calculations, the finite element mesh model of the turbine disc is create by using a density transition approach. In the stress concentration regions, such as the firtree rim region, the assembly holes and the hub region, 10- node tetrahedral units were used to divide the dense grids. In the turbine disc that has regular structure and low stress gradient, 8-node hexahedral units were used to divide the sparse grids. During the operation of an aero-engine, the turbine disc is subjected to a mixed load: centrifugal forces, thermal stresses, aerodynamic forces and vibratory stresses", " Then the finite element analysis results of the 1/6 turbine disc segment under different working cycles are shown in Figs. 5-7 below. As Figs. 5-7 show, for the 1/6 turbine disc segment, the maximum stress and strain concentration occurs on the assembly hole. The stress and strain concentration locations can determine the lifetime of 1/6 turbine disc segment so the most critical high stress region of the 1/6 turbine disc segment is located in the assembly hole. Similarly, for the three-dimensional model in Fig. 4, the simplified 1/49 turbine disc segment is also subjected to a combined load of centrifugal forces and thermal stresses. The centrifugal forces were loaded on the three-dimensional model in the form of rotational speed. And the thermal stresses were loaded into the three-dimensional model by inputting the temperature data with ANSYS parametric design language. Besides, the contact between turbine disc and blade was considered and a friction coefficient of 0.15 was defined. The finite element analysis results of the 1/49 turbine disc segment under different working cycles are shown in Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002968_j.mechmachtheory.2020.104186-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002968_j.mechmachtheory.2020.104186-Figure11-1.png", "caption": "Fig. 11. Exoskeleton kinematic structure of an individual obtained with the EM approach. The two symmetric arms are reported.", "texts": [ " The remainder of the \u03b5 parameters has then been defined by the EM algorithm. The orientation parameter \u03b52 and the offset parameters \u03b53 , 4 have been permuted and 64 exoskeleton structures have have been obtained. However, only 12 have respected the evolution criterion 1 , six of which have also respected the criterion 2 , thus causing the algorithm to terminate. The final generation of 6 exoskeletons fulfils the requirement \u03c61 \u22126 . The DH parameters ( Table 2 ) and the resulting kinematic structure of one of these final solutions is shown in Fig. 11 as a prime example. The article presents a new method for the structural design of kinematic structures that relies on both the graph theory and the theory of linear transformation, combining two classical approaches in order to overcome their main limitations, i.e. the planarity assumption of the former [19] , and the high computational complexity of the latter [14,15,33,34] . The proposed method has been applied to a specific scenario, i.e. the type synthesis of a shoulder exoskeleton, chosen in light of the challenging requirements of kinematic compatibility that the human component introduces in the problem of exoskeleton\u2019s structural synthesis [5,6,25,28,29,35,75\u201377] and the actual need of exoskeleton devices for the reduction of musculo-skeletal disorders associated with particular working environment and conditions [37,38,78] " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001851_s12206-016-0614-7-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001851_s12206-016-0614-7-Figure11-1.png", "caption": "Fig. 11. Schematic diagram of the apparatus for the tests.", "texts": [ " Therefore, all parameters about Coulomb friction in this study were found through trial and error. Table 1 shows the specific value of parameters of Coulomb friction. The modified model was developed through the above described process. Fig. 9 shows the multi-body dynamic model of the taper leaf spring of tandem suspension. A series of tests were conducted to verify the modified model. The taper leaf spring of tandem suspension used for the tests, as shown in Fig. 10, had four leaves, 90-mm width, and non-uniform thickness. Fig. 11 is a schematic diagram of the apparatus for the tests. Fig. 11 shows that the hydraulic actuator, which is equipped with high-precision sensor in the head and controlled by computer, was used to apply the sinusoidal alternating load to the taper leaf spring of tandem suspension. The excitation frequency was controlled by the vertical velocity of the hydraulic actuator. The different sinusoidal alternating load was applied at the middle of the taper leaf spring of tandem suspension. The tests were divided into two groups. The first group started at an unloaded state with the excitation frequency of 1/30 Hz and the amplitude of 70 mm to acquire quasi-static behavior" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001235_j.acme.2014.11.003-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001235_j.acme.2014.11.003-Figure8-1.png", "caption": "Fig. 8 \u2013 The decoupled deformation of", "texts": [ " There is also a possibility of direct transmission of the vibration energy from gears through the oil in radial gap to the pump housing. From investigations in [2] it is known that the transmission of the vibrational energy through this way is very small. Please cite this article in press as: W. Fiebig, M. Korzyb, Vibration a Mechanical Engineering (2015), http://dx.doi.org/10.1016/j.acme.2014.11.0 The gears are deflected toward the suction side due to pressure loads. These deflections are dependent on the bearing clearance and the deflection of the wheels. In Fig. 8 the displacements of the shaft center xw, yw from the static equilibrium position have been schematically shown. Absolute displacement of the shaft center OstatOw is resulting from the shift due to the flexibility OstatOst of the supporting housing, from the displacement due to the elasticity of the oil film in the bearings OstOz and shaft deflection OzOw. The stiffness of the sliding bearings at small oscillations around the static equilibrium position has been linearized with four stiffness and damping coefficients cik, bik (i, k = 1, 2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure21.13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure21.13-1.png", "caption": "Figure 21.13 Symbols for system of projection", "texts": [ "5 Type of line Description of line Application Thick, continuous Visible outlines and edges Thin, continuous Dimension and leader lines; projection lines; hatching Thin, short dashes Hidden outlines and edges Thin, chain Centre lines Chain, (thick at ends and at changes of direction, thin elsewhere) Cutting planes of sections External thread Round shaft Rectangular block Interrupted views Holes on a linear pitch Internal thread Square on a shaft Figure 21.14 Conventional representation All lines should be uniformly black, dense and bold. Types of line are shown in Table 21.5. The system of projection used on a drawing should be shown by the appropriate symbol as shown in Fig. 21.13. Standard parts which are likely to be drawn many times result in unnecessary waste of time and All dimensions necessary for the manufacture of a part should be shown on the drawing and should appear once only. It should not be necessary for a dimension to be calculated or for the drawing to be scaled. Dimensions should be placed outside the outline of the view wherever possible. Projection lines are drawn from points or lines on the view and the dimension line placed between them. Dimension lines and projection lines are thin continuous lines" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure7.17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure7.17-1.png", "caption": "Figure 7.17 Bench grinder", "texts": [ " Because of this, its use requires stringent safety precautions. The off-hand grinding machine is basically an electric motor having a spindle at each end, each carrying a grinding wheel, also referred to as an abrasive wheel. This arrangement allows a coarse wheel to be mounted at one end and a fine wheel at the other. All rough grinding is carried out using the coarse wheel, leaving the finishing operations to be done on the fine wheel. These machines may be mounted on a bench, when they are often referred to as bench grinders (Fig. 7.17), or on a floor-mounted pedestal and referred to as pedestal grinders. Bench grinders should be securely anchored to a stout bench and pedestal grinders should be heavily built and securely bolted on good foundations. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 7 Cutting tools and cutting fluids 110 7 gap at a minimum. Adjustments of the work rest should only be done when the grinding wheel is stationary and the machine isolation switch is in the \u2018off \u2019 position" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure6-1.png", "caption": "Fig. 6. Definition of the screw axes a\u03021 and a\u03022.", "texts": [ " Algorithm for the numerical synthesis The algorithm for the numerical synthesis of overconstrained mechanisms is exemplarily described for a 6H mechanism, thus n = 6. In general the 36 unknowns for the algorithm are the screw coordinates ai1, ai2, ai3, ai4, ai5, ai6, i = 1, . . . , 6. Again the number of unknowns can be reduced by special choice of the reference frame. Without loss of generality the axis of the joint screw a\u03021 coincides with the x-axis of the base frame, a\u03021 = [ 1 0 0 h1 0 0 ]T . (65) Again without loss of generality the axis of the second joint screw can be defined by the geometric parameters , a21, a23 according to Fig. 6 as a\u03022 = [ a21 0 a23 a23 + h2a21 0 \u2212 a21 + h2a23 ]T . (66) The unknowns are now given by \u2022 the pitch of the first joint screw h1 (1 unknown) \u2022 the pose of the second joint screw, expressed by a21, a23 and (3 unknowns) \u2022 the screw coordinates ai1, ai2, ai3, ai4, ai5, ai6, i = 3, . . . , 6 (24 unknowns). These altogether 28 unknowns are summarized in the vector x. The aim of the synthesis algorithm is to find numerical values x\u2217 and consequently coordinates for the joint screws a\u0302\u2217 i = a\u0302i(x\u2217), i = 1, " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002071_asjc.1315-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002071_asjc.1315-Figure1-1.png", "caption": "Figure 1. Two-dimensional airfoil model with wing flap.", "texts": [ " A nonlinear fault-tolerant controller is proposed for airfoil flutter using the SMC method. The paper is organized as follows. \u00a9 2016 Chinese A Section II presents the dynamic model of wing flutter. The adaptive sliding mode fault-tolerant controller is presented in section III. In section IV, Lyapunov stability analysis is carried out. Section V gives the results of numerical simulations, and section VI draws some conclusions. In this section, a flutter model for a two-dimensional wing, including cubic hard spring nonlinearity is established. As shown in Fig. 1, a two-degree-of-freedom wing model is considered. The plunge deflection is denoted by h, positive in the downward direction; \u03b8 is the pitch angle about the elastic axis, positive nose up; the span of the wing is b; the chord length is c; Q and P are the wing aerodynamic center and elastic axis, respectively; the distance from the leading edge to the elastic axis is xf; the distance from the aerodynamic center to the elastic axis is ec; e is the eccentricity; and \u03b2 is the deflection angle of control surface", " Differentiating (21) and considering (7)\u2013(17) result in _V \u00bc sCs _x- _\u03bb 1 \u03b5\u03c3 \u03b1 e\u03bc _b\u03bc \u00bc sfCs eAx Cs eAx\u00fe Cs ef -CsG\u03bb ks\u00fe \u03c2 sgn s\u00f0 \u00de\u00bd \u00fe CseB\u0394u\u00fe CseB \u03c3-1\u00f0 \u00de\u03c7\u03c5 \u00feCs ef -CsG\u03bb CseB\u0394ug 1 \u03b5\u03c3 \u03b1 e\u03bc _b\u03bc \u00bc s ks\u00fe \u03c2 sgn s\u00f0 \u00de\u00bd \u00fe CseB \u03c3-1\u00f0 \u00de\u03c7\u03c5 n o 1 \u03b5\u03c3 \u03b1 e\u03bc _b\u03bc \u2264 ks2 \u03c2 sj j \u00fes \u03c3-1\u00f0 \u00de\u03c7CseB\u03c5 1 \u03b5\u03c3 \u03b1 e\u03bc _b\u03bc (22) In particular, from (16) and the controller (11), we have s \u03c3-1\u00f0 \u00de\u03c7CseB\u03c5 \u00bc s \u03c3-1\u00f0 \u00de\u03c7 ( Cs eAx\u00fe Cs ef -CsG\u03bb ks\u00fe \u03c2 sgn s\u00f0 \u00de\u00bd g \u2264 s\u03b5\u03c3 Cs eAx\u00fe Cs ef -CsG\u03bb\u00fe ks \u00fe \u03c2 \u2264 sj j\u03b5\u03c3b\u03bc\u0393 (23) Imposing the bound (23) and\u03bc \u00bc 1 1 \u03b5\u03c3 , (22) can be fur- ther evaluated as _V\u2264 ks2 \u03c2 sj j \u00fe sj j\u03b5\u03c3b\u03bc\u0393 1 \u03b5\u03c3 \u03b1 e\u03bc _b\u03bc \u2264 ks2 \u00fe sj j\u0393 \u03b5\u03c3b\u03bc-1\u00fe b\u03bc\u00f0 \u00de 1 \u03b5\u03c3 \u03b1 e\u03bc _b\u03bc \u2264 ks2 < 0 (24) utomatic Control Society and John Wiley & Sons Australia, Ltd. \u00a9 2016 Chinese Automatic Control Society and John Wiley & Sons Au Based on Lyapunov stability theory, the above proof can prove that the closed-loop system is stable in the sense of Lyapunov. In this section, numerical simulations are carried out to demonstrate the validity of the proposed method in this paper. The two-dimensional airfoil shown in Fig. 1 is considered, and the structural parameters of the airfoil are displayed in Table I. Before the fault tolerant control for the airfoil, we use the classical V-g method to calculate the critical flutter speed of the airfoil. The V-g method only requires a straightforward complex eigenvalues analysis to be done for all values of reduced frequency. Flutter speed is located at the point where the value of the damping stralia, Ltd. becomes positive [14]. Fig. 2 shows the relationship between the wind speed and the artificial damping" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001279_s0263574714000149-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001279_s0263574714000149-Figure3-1.png", "caption": "Fig. 3. Mechanism with closed loops and the corresponding representational graph.", "texts": [ " how many equations correspond to joints and how many correspond to the driving motions. It also provides information regarding the dimensions and structure of the DAE system for the dynamic formulation, i.e. how many differential and how many algebraic equations. Nevertheless, this structural model does not involve any information regarding the number of closed loops and their connectivity. According to this model, a mechanism is seen as a set of serial kinematic chains inter-connected through relative constraints. The mechanism structure is usually represented by a graph (Fig. 3), which enables an easy identification of loops. This model9 is very popular, especially in the kinematic and dynamic analysis of robots and other controlled mechanical systems as it facilitates the direct control of joint motion parameters. Within this model, the KCB is thus an open chain of bodies serially interconnected through joints and represents the structural primitive building block of a mechanism. Its dof is variable with the number and type of individual joints integrated by the KCB (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003768_adem.202100265-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003768_adem.202100265-Figure1-1.png", "caption": "Figure 1. Illustration of compression and SEM samples extracted from different specimen heights.", "texts": [ " To evaluate the effect of platform temperature on the microstructure and mechanical property of as-built Ti-55531, we investigated two types of platform temperatures: one was at 200 C, which was commonly used in fabricating SLM Ti alloys (named S200), and the other one was 700 C (named S700). The microstructure characterizations and mechanical property evaluations were performed at different deposition heights, i.e., bottom region (Z\u00bc 5mm), middle region (Z\u00bc 30mm), and top region (Z\u00bc 58mm), as schematically shown in Figure 1. X-ray diffraction (XRD; D8 Advance, Bruker; step size of 0.01 ), field emission scanning electron microscope (SEM; FEI-Inspect F50), and transmission electron microscopy (TEM; FEI-Tecnai G2 F20) were used to analyze the phase and microstructure of as-fabricated Ti-55531 at different substrate temperatures. SEM and nanoindentation samples were prepared by using a Buehler AutoMet 250 polisher. SEM samples were etched by Kroll\u2019s agent (1 vol% HF \u00fe 4 vol% HNO3\u00fe 95 vol% H2O). TEM foils were prepared by ion beam milling (Gatan-PIPS II 695)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001483_0954406215612814-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001483_0954406215612814-Figure2-1.png", "caption": "Figure 2. Generation of two conjugate involute tooth profiles via the Camus theorem.", "texts": [ " The generic positions of \"0l and \"0l are shown in Figure 1(b), along with the positions P0l and P0l of the tracing point P. At the starting position, the generated profiles and s are tangent to each other at the contact point P; they remain constantly in contact during their relative motion, represented by the purerolling of l on l for a given pair of conjugate profiles and s, as shown in Figure 1(c). Camus\u2019 theorem can be also applied to the generation of the involute tooth profiles of circular and non-circular gears in the form of the well-known rack-cutter method. In fact, referring to Figure 2, when the auxiliary centrode \" becomes the tangent line t to both centrodes l and l, and its attached curve becomes also a straight line, a pair of involute conjugate profiles can be generated as envelope of , while their normal line N at their contact point P envelops the corresponding base curves, which are at UNIV CALIFORNIA SAN DIEGO on December 14, 2015pic.sagepub.comDownloaded from also the loci of the centers of curvature of each involute tooth profile. The centers of curvature and s of the involute tooth profiles and s are also shown in Figure 2, as obtained by applying the Euler\u2013Savary equation for envelopes via a graphical construction. In particular, the center of curvature s of profile s coincides with the contact point P, which means that the involute profile s has a cusp at P. In fact, the line attached to the auxiliary centrode \" (line) passes in this case through the center of curvature Ol of the centrode l. This feature will be clarified in the Euler\u2013Savary equation for envelopes and Aronhold theorems and the return circle sections based on the Euler\u2013Savary equation for envelopes and Aronhold\u2019s first theorem, which takes into account the properties of the return circle", " Consequently, both centers of curvature and s can be joined through the coupler link s to produce a linkage, which is instantaneously equivalent, up to their accelerations, to the corresponding mechanism with the higher-kinematic pair represented by the synthesized conjugate profiles and s. This is the basic concept to convert the conjugate tooth profiles and s into the instantaneously equivalent linkage Ol s Ol, whose coupler link is s. The Euler\u2013Savary equation for envelopes can be also applied to the particular case of involute gears, whose graphical construction is shown in Figure 2. In fact, the center of curvature 1 of the straight line should be joined with that of the auxiliary centrode \" up to intersect, at point H, the normal line to N at point P0. Consequently, the improper points 1 and O\"1 must be aligned along the line at infinity, which means that H is also an improper point H1. Thus, the centers of curvature and s are obtained as intersections with N , of two parallel lines passing through Ol and Ol, respectively. Aronhold theorems and the return circle The return circleR is the locus of the centers of curvature of all points at infinity of a moving plane, whose equation can be readily derived from the Euler\u2013 Savary equation 1 P0 1 P0M cos \u00bc 1 rl 1 r l \u00bc 1 \u00f02\u00de where P0M and are the polar coordinates of a generic moving point M with respect to the canonical frame (P0, t, n) shown in Figure 4, and is the corresponding center of curvature" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001620_s12206-015-0908-1-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001620_s12206-015-0908-1-Figure10-1.png", "caption": "Fig. 10. Static stress and strain state under cruise cycle.", "texts": [], "surrounding_texts": [ "During the operation of an aero-engine, the turbine disc is subjected to a mixed load: centrifugal forces, thermal stresses, aerodynamic forces and vibratory stresses. Of course, high speed results in large centrifugal forces and high thermal gradients result in thermal stresses. Among them, the aerodynamic forces and vibratory stresses have little effect on the static strength of the turbine disc. Therefore, when analyzing the turbine disc with finite element method, the centrifugal forces and thermal stresses are the main consideration. The speed spectrum of the turbine disc is determined by the flight mission, and it consists of three parts [11]: low frequency cycle, full throttle cycle and cruise cycle. Any speed spectrum can be considered as a combination of these three basic cycles. The speed spectrum of the turbine disc is shown in Table 3. The temperature spectrum is derived based on the measurement data. In this study, the temperature spectrum of the turbine disc was loaded on the three-dimensional model by ANSYS parametric design language. For each basic cycle mentioned above, there are 100000 temperature data points of the turbine disc. Table 4 shows part of the temperature data points under full throttle cycle, where X, Y and Z represent the coordinate value of a point of the three-dimensional model." ] }, { "image_filename": "designv11_22_0000486_s11771-019-4180-x-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000486_s11771-019-4180-x-Figure1-1.png", "caption": "Figure 1 Meshed pinion and gear three-tooth model", "texts": [ " The gear 3D CAD models were constructed in Autodesk Inventor and then sliced to leave only a pair of three J. Cent. South Univ. (2019) 26: 2368\u22122378 2371 toothed gear segment. The static stress analysis was conducted in ANSYS Workbench. The gear volumes were mapped-meshed by sweeping it with SOLID185, a 3D finite element without mid-side nodes. The contact area was assigned with the contact elements CONTA174 and TARGE170. The gear involute curve was divided into 64 divisions to increase mesh density around the contact area as shown in Figure 1. The meshed model has 144300 elements and 165165 nodes with good mesh size. The mesh convergence test was performed to determine the optimum mesh size. The load and boundary conditions were applied as shown in Figure 2. The left and right sides of the gear and hub were fixed in all directions. The left and right sides of pinion were fixed tangentially while allowed to move freely in the axial and radial directions. Meanwhile, the pinion hub was fixed in the axial and radial directions and allowed to freely rotate tangentially" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003465_j.precisioneng.2021.03.011-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003465_j.precisioneng.2021.03.011-Figure1-1.png", "caption": "Fig. 1. Configuration of the USM. (a) Principles of operation for the USM. (b) Structure. (c) Stator. (d) Flexible shaft and rotor.", "texts": [ " In continuous mode, the speed of the motor was controlled by PID control. In step mode, the step characteristics of the motor was analyzed and tested under different load torques and different lengths of driving voltage to find the parameters which could achieve minimum step angle. Finally, the positioning performances of USM were tested. A USM is mainly composed of a stator, a rotor, a shaft, bearings and shell. By electrical excitation of a piezoceramics ring, which is bonded to the lower surface of the stator, a bending-wave of the shape illustrated in Fig. 1 (a) is excited in the stator. The structure of stator is plate-type. Thus, the traveling wave causes points on the stator surface to run through elliptic trajectories. The teeth on stator surface are used to enlarge vibration amplitude. With the rotor pressed against the stator, frictional force generated from the contact zones of stator and rotor drives the rotor. For the high-precision positioning of the USM, the structure should be improved to make the contact of stator and rotor more uniform and reduce the clearance in the shafting. The contact of stator and rotor in USMs is surface to surface, so the tilting of the contact surface is inevitable. The tilting is acceptable in normal conditions, but it would affect the stability of micro-step performances of the motor. In this research, a flexible shaft and central preload method was proposed to improve the contact. Fig. 1 (b) shows the structure of the USM. The motor used in this paper consists of a stator, a rotor, a front shell, a flexible shaft, a rear shell, a preload nut and two bearings. Compared with the normal USMs, a preload nut on shaft is used to apply the pressure between stator and rotor instead of shell bolts. The diameter of the stator is 60 mm. This central preload method could make the pressure more uniform. A flexible hinge was placed on the shaft between the stator and rotor. The deformation of the flexible shaft can compensate for the uneven contact of stator and rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003718_j.mechmachtheory.2021.104425-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003718_j.mechmachtheory.2021.104425-Figure1-1.png", "caption": "Fig. 1. Schematic of involute helical surface S 1 .", "texts": [ " is the involute helical surface. Thus, the rigorous proof is made to prove whether the worm tooth surface with constant helix parameter ground by a planar grinding wheel is an involute one using the uniqueness theorem for surfaces [33] . The mapping and rigid body transformation relations between these two surfaces are also studied. This proof complements the machining principle of the machining method proposed by D. Brown Co. for the involute worm ground with the planar grinding wheel. As shown in Fig. 1 , a unit orthogonal rotation frame \u03c31 { O 1 ; i 1 , j 1 , k 1 } is fixed on the base cylinder of the right-hand involute helical surface and used to reflect the current position of the worm. The radius of the base cylinder for the involute helical surface is denoted as r b . The unit basis vector k 1 lies along the axial line of the base cylinder of the involute helical surface. The unit basis vectors i 1 and j 1 are in the cross-section of the base cylinder. The symbol S 1 represents the involute helical surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003601_tmag.2021.3081799-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003601_tmag.2021.3081799-Figure2-1.png", "caption": "Fig. 2. Sectional view and trimetric view of 18/12/17 FM-TFLG. (a) Sectional view. (b) Trimetric view.", "texts": [ " Generally speaking, the 12-stator-pole machine with 11- translator-pole translator has better performance than 12/10, 12/13 and 12/14 counterparts [11]. Such as, higher power density, smaller thrust ripple and better sinusoidity of back EMF. In order to study whether higher modulation ratio can obtain better performance, such as higher power density, better sinusoidity of back EMF and higher efficiency, 12/12/11 and 18/12/17 FM-TFLGs are analyzed and compared in this paper, the translator numbers of both topologies are one less than the PM numbers of these two machines. The configuration of 18/12/17 FM-TFLG is shown in Fig. 2. The permanent magnet number, slot number, translator number of the generator are 18, 12, 17, respectively. In the proposed FM-TFLG, armature winding pole pair awp , translator pole number tN , and PM pole pair PMp match [10]: t PM awN p p (1) The magnetomotive force (MMF) equation is expressed as follow: 0= /e r m rF B h (2) where rB is the residual magnetic flux density, mh is PM thickness, 0 is vacuum permeability, r is the relative permeability. In the proposed machine, the three-phase symmetric sinusoidal currents can be expressed as: 2 sin 2 sin 2 / 3 2 sin 2 / 3 A rms e B rms e C rms e i I t i I t i I t (3) where rmsI and e are phase root mean square current and electric angular speed respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003178_j.jsv.2015.06.037-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003178_j.jsv.2015.06.037-Figure1-1.png", "caption": "Fig. 1. The layout of an impeller\u2013shaft-bearing system.", "texts": [ " For structures rotating at a constant angular velocity \u03a9, the governing equations of motion in rotating frames can be expressed as [25] Ms \u20acus\u00fe Cs s\u00feCs b\u00fe\u03a9Cs cor _us\u00fe Ks s\u00feKs b \u03a92Ks d us \u00bc fsint\u00fefsext (1) where superscript \u201cs\u201d is a label representing structures; Ms is the mass matrix; Ks s is the elastic stiffness matrix; Cs b and Ks b denote the damping and stiffness matrices due to bearings; \u03a9Cs cor is the skew-symmetric Coriolis matrix in rotating frame; Cs s represents other damping factors; \u03a92Ks d is the spin softening matrix due to the rotation of structure; us is the vector of nodal displacement; fsint is the interface reactions on the crack surfaces; and fsext denotes external forces. The rotor studied in this paper is shown in Fig. 1. The rotor is a rotating assembly of a centrifugal compressor and is composed of an impeller, a shaft and two journal bearings. The impeller is manufactured by welding cover, blades and disk components. Fig. 2 shows the sector finite element model of the impeller\u2013shaft assembly. As is illustrated in the figure, the crack locates at the weld toe on the cover sides of blade. In the modeling process of the rotating assembly, the single-sector model is partitioned into two substructures, namely cover sector and rest part" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000868_gt2015-43971-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000868_gt2015-43971-Figure6-1.png", "caption": "Figure 6: A profiled substrate was pre-machined using a CNC lathe.", "texts": [ " Building Impellers Laser consolidation is a material addition process that can directly build functional features on an existing component to form integrated structure without the need of welding or brazing. For example, LC can be used to build net-shape functional components on pre-machined substrate. An impeller shape (Figure 5) was selected to demonstrate the capability of LC process using a 5-axis CNC motion system. The impeller has a diameter of about 77 mm and height of about 26 mm. There are 9 long blades and 9 short blades uniformly distributed. A profiled substrate was pre-machined using a CNC lathe (Figure 6) and was mounted on a 5-axis CNC motion system using a simply designed fixture. Using a 5-axis CNC motion system to deal with large tiltrotation movement, LC was successfully conducted to build blades on the pre-machined substrate to form an integrated impeller. Figure 7 shows an IN-718 impeller (left) with its blades directly built up on a pre-machined substrate using LC process. After sand blasting, laser-consolidated blades and premachined substrate show consistent surface finish (Figure 7, right side)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000580_epe.2019.8915161-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000580_epe.2019.8915161-Figure1-1.png", "caption": "Fig. 1. (a) 3-D simulation domain of the InAs/Si lateral NW TFET along with the plane of vertical cross section. Inset: inclination of the \u3008111\u3009 InAs/Si interface to the channel direction, where gate oxide and InAs regions were made transparent. (b) 2-D cross section of the device with various lengths. (c) Nonlocal mesh generated to model TAT at the InAs/oxide and InAs/Si interfaces.", "texts": [ " An intrinsic underlap region between i-Si channel and p+-Si drain is present in FF44, while no such underlap exists in FF41. The InAs region is n+-doped with a concentration of 2 \u00d7 1018 cm\u22123. The section of Si in the channel and the underlap region are intrinsic. The overlap between the InAs source with the gate is 270 nm long while that between Si and the gate is 730 nm long. The simulated underlap region in FF44 has an extension of 100 nm. The gate oxide consists of HfO2/Al2O3 with an effective oxide thickness (EOT) of 1.75 nm. A vertical cross section of the device at the midpoint along the width of the NW [Fig. 1(a)] was simulated using the commercial simulator S-Device [9]. Since the channel orientation is along the \u3008100\u3009-direction while the InAs/Si interface is a (111)-plane, the InAs/Si interface in the cross section is slanted, making an angle of 450 with the buried oxide. It must be noted that the device does neither possess a symmetry axis nor a plane of symmetry. A vertical cross section was selected for 2-D simulations to lower the computational burden, but some 3-D simulations were also done for validation purposes", " At the InAs/Si interface, the energetic distribution of traps was assumed to have Gaussian shape with a peak concentration of 6 \u00d7 1013 cm\u22122eV\u22121, a full-width-half-minimum (FWHM) of 0.11 eV, and the peak position at the VB edge of InAs (see Section III for a more detailed discussion). For the simulation of FF41 TFETs, the trap density at the Si/oxide interface was assumed to be uniform throughout the bandgap of Si. Its value was set to 4 \u00d71012 cm\u22122eV\u22121 after fitting the experimental data. The FF44 TFET [see Fig. 1(b)] was simulated by including the traps at the InAs/Si and InAs/oxide interfaces and using the model parameters outlined in Section II. A comparison of simulated and measured transfer characteristics is shown in Fig. 3. The used model parameters are mostly experimental data; only Huang\u2013Rhys factor, trap interaction volume, and trap density at the InAs/Si interface, which are related to the TAT model, have been fitted to obtain a good match to the experimental data. It is observed that the variation of the Huang\u2013Rhys factor has only a small effect on the transfer characteristics" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure4.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure4.9-1.png", "caption": "Figure 4.9 Folding operations", "texts": [ " D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 4 Sheet-metal operations 57 to allow for various thicknesses of material and can be made up in sections known as fingers to accommodate a previous fold. Slots between the fingers allow a previous fold not to interfere with further folds, as in the case of a box section where four sides have to be folded. The front folding beam, pivoted at each end, is operated by a handle which folds the metal past the clamping blade, Fig. 4.9. With the use of simple tools, the fly press can also be used for bending small components, Fig. 4.7. The top tool is fixed to the moving part and the bottom tool, correctly positioned under the top tool, is fixed to the table of the press. Metal bent in this way will spring back slightly, and to allow for this the angle of the tool is made less than 90\u00b0. In the case of mild steel, an angle of 88\u00b0 is sufficient for the component to spring back to 90\u00b0. Figure 4.6 Punch and die in fly press Punch in fly-press spindle Piece removed Metal sheet Die clamped to fly-press table Former in fly-press spindle Metal sheet Block clamped to fly-press table Figure 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure21.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure21.2-1.png", "caption": "Figure 21.2 Exploded view of hand riveting tool", "texts": [ " Guides to the use of BS 8888 are available for schools and colleges and for further and higher education as PP 8888 parts 1 and 2. Adopting a system of standard specification, practices and design results in a number of advantages: XX reduction in cost of product; D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 21 Drawings, specifications and\u00a0data 21 306 dimensions. They are used for assembly purposes and in service or repair manuals where reference numbers of the parts and the way in which they fit together are shown. Figure 21.2 shows an exploded view of a hand riveting tool, listing the spare parts, their identification numbers and their relative position within the finished product. XX overheads are reduced. Except for written notes, technical drawings have no language barriers. They provide the universal language for design, for the craftsman and the technician in manufacture, assembly and maintenance, for the sales team as an aid to selling and for the customer before buying or indeed servicing after purchase. Many different methods are used to communicate technical information" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003528_s00170-021-06664-9-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003528_s00170-021-06664-9-Figure2-1.png", "caption": "Fig. 2 Automatic WCS calibration using a single camera with different positions. a The robot controls the single camera\u2019s movements, and images of an object acquired by the camera at different viewing positions are used to restore its 3D location. b The sketch map of camera vision calibration and relevant coordinate systems", "texts": [ " The proposed system achieves a performance that is comparable with implementations of the previous manual calibration method called human vision calibration (HVC) and has a more accurate repairing appearance. The rest of the paper is organized as follows: Section 2 provides an overview of the camera vision calibration method, followed by image processing and visual calibration in the 3D reconstruction. Section 3 presents the results of real-world experiments and the sources of systematic errors. Section 4 closes with the conclusions. Calibrating WCS is to calculate the relative pose between the WCS (denoted by the letter {W}) and the robot base coordinate system (denoted by the letter {B}). Figure 2 illustrates the idea. The robot is equipped with cameras that provide information about the parts to manufacture. The camera\u2019s movements are precisely controlled by the robot, and the images of the feature points P1, P2, P3 are acquired by the camera at different viewing positions. Structure-from-Motion (SfM) [28, 29] is used to furnish those feature point 3D coordinates under the camera coordinates (denoted by the letter {C}). The coordinates need to be transformed into the robot base coordinates {B} by the transformation matrixes between the coordinate systems {W}, {C}, {tool}, and {B} shown in Fig. 2b. Finally, we can use those non-collinear points to establish the workpiece coordinates {W}, which now is under the coordinate system {B}. In Fig. 2a, the vector Vx=P2-P1 is the X-axis of the frame {W}, and the pedal (O) of P3 and V1 is the origin. The Y-axis is the vector Vy=P3-O, and the Z-axis is the vector Vz =Vx \u00d7Vy. The calibration process is shown in Algorithm 1 . The shape and the high temperature of the moldmake it harder to install the calibration block. We drill small holes (about 1 mm in diameter) as feature points on the mold\u2019s surface instead. Small feature points provide better positioning accuracy in off-line programming software but make image processing more difficult in CVC" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003241_access.2021.3050098-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003241_access.2021.3050098-Figure1-1.png", "caption": "FIGURE 1. The carbon brush slip-ring system abrasion test platform using a salt spray tester.", "texts": [ " Based on certain amount of experimental data, the contact current, contact pressure, salt spray concentration and experimental temperature were introduced into the wear model, the unknown model parameters were identified by using the least square method on the test data. This abrasion prediction model can reflect different contact currents, contact pressures, salt spray concentrations, and experimental temperatures. The validity of the model has been verified through experiments. II. EXPERIMENT AND FACILITIES A. EXPERIMENTAL FACILITIES We studied abrasion on the Doubly fed machine carbon brush and slip-ring system using a salt spray tester research and experiment platform. Fig. 1 illustrates the brush slipring system abrasion test platform using a salt spray tester. The test platform mainly comprised a salt spray test box, asynchronous motor, frequency converter, voltage regulator, grid system, carbon brush and slip-ring system, and sensor. The salt spray test box in the testing machine could alter the experimental corrosive environment using a 1%-7%NaCl solution. The combination of an asynchronous motor and frequency converter could convert the carbon brush slip-ring system in a revolving speed range of 0-1500 r/min" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000354_iemdc.2019.8785369-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000354_iemdc.2019.8785369-Figure1-1.png", "caption": "Fig. 1. NovaMAX motor (a) transverse-section view with red arrows illustrating flux direction, (b) cross-section view of one rotor with red arrows illustrating flux direction, and (c) view with one rotor removed.", "texts": [ " Keywords\u2014 Axial flux, conical rotor, dual rotor, efficiency, flux concentration, grain oriented electrical steel, losses, motor, permanent magnet, soft magnetic composite I. INTRODUCTION The US Department of Energy (DOE) Advanced Manufacturing Office\u2019s Next Generation Electric Machines: Enabling Technologies program seeks to develop ways to employ high performing materials to improve the efficiency of motors without depending on large quantities of rare earth materials [1]. Regal Beloit\u2019s NovaMAX motor incorporates several innovations to achieve very high efficiencies. As shown in Fig. 1, the NovaMAX motor is a dual-rotor permanent magnet (PM) motor with conically shaped air gaps. Grain oriented electrical steel (GOES) provides higher permeability, higher saturation flux density, and lower core losses than nonoriented electrical steel when the flux is primarily in the direction of the grain orientation. Thus, GOES is frequently used in transformer cores; however, because most motors employ a rotating magnetic field, nonoriented steel is used in the vast majority of motors. Nonetheless, GOES has been applied to motors in a few studies [2]-[6], but the manufacturing complexity of the motor is often increased to accommodate the GOES [2]-[5]. However, as in [6], the dual rotor axial topology results in the individual stator teeth only being exposed to a pulsating axially directed flux, rather than a rotating flux, so the teeth in the NovaMAX motor are well suited for using GOES. Additionally, the NovaMax motor employs a flux concentrating rotor topology. As illustrated in Fig. 1(b), flux This work was supported by the US Department of Energy under award number DE-EE0007875. 978-1-5386-9350-6/19/$31.00 \u00a92019 IEEE 1067 from PMs in multiple directions is concentrated in soft magnetic composite (SMC) pole pieces in each rotor before crossing the air gap to the stator. Furthermore, the conical shape of the air gap results in a larger surface area relative to an axial flux motor with the same diameter. This increased air gap surface area reduces the reluctance of the flux path, resulting in an increase in flux traveling from the rotor to the stator" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001992_icma.2016.7558710-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001992_icma.2016.7558710-Figure9-1.png", "caption": "Fig. 9. Foot contact configuratoins during swing phase.", "texts": [ " 3) Heel Clearance Pattern: After the ankle position is determined, we can determine the ankle angle during swing phase by using heel clearance pattern. The heel clearance means the interval between the ground to the heel. We define the constraint for generating the heel clearance pattern to meet that the foot is unable to contact the ground and the motion can be kinematically admissible during the swing phase as follows: zA \u2212 La < zH < zA + La, zA > Lb (6) case A : zHb < zH < zA + La, zA Lb case B : { zA \u2212 La < zH < zHb, La < zA Lb 0 < zH < zHb, zA La . (7) When the ankle height (zA) is lower then Lb as shown in Fig. 9, there are two cases of foot contact configuration. A boundary of zH (zHb) is the height when the toe contact on the ground and we can define zHb as follows: \u03b2 = arccos ( LFF Lb ) , \u03b3 = arcsin ( zA Lb ) (8) \u03b1 : { \u03c0 \u2212 (\u03b2 + \u03b3), case A \u03b3 \u2212 \u03b2, case B (9) We know the range of zH by using the constraint and zHb and then we can define zH trajectory with the heel clearance from human walking (Fig. 8 (b)). We also generate zH by using piecewise polynomial interpolation. After zH is defined, we can define the x position of the heel (xH ) using kinematic configuration (11) and complete the swing phase pattern generation finally" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000588_9783527822201.ch15-Figure15.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000588_9783527822201.ch15-Figure15.5-1.png", "caption": "Figure 15.5 Schematic representation of the three SSBAs. The robotic root is composed of a module with the three motors driving the springs, a three-shaft module where the springs slide in and a sensorized tip for controlling behavior.", "texts": [ " Natural roots bend and follow stimuli by means of differential elongations of the cells at the apical level, a mechanism that improves the penetration capabilities because it decreases the lateral friction. To achieve the same functionalities, a novel soft spring-based actuator (SSBA) system has been proposed in [13], making use of helical springs to transmit the motor power in a compliant manner. This novel pan-tilt mechanism is able to bend in each direction, by modulating the length of the springs that connect two planes, and at the same time is flexible and compliant thanks to the use of soft springs. The mechanism is based on three helical springs (Figure 15.5), located at 120\u2218 to each other, used as screws in a nut\u2013screw mechanism that translates the rotational motion in a linear displacement. This soft linear actuator can transmit the motion even after buckling (bending of the spring) and thanks to its compliance, a differential elongation between the three springs results in a bending of the system. Each SSBA can elongate and retract by changing the rotation of the motor connected to it. The robotic root elongates straightly when all three SSBAs elongate at the same speed, and performs the bending when the elongation velocity is different (differential elongation)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure10.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure10.9-1.png", "caption": "Figure 10.9 Dressing the grinding wheel using a diamond", "texts": [ " The opposite is true when grinding a soft material: the grains do not blunt so readily and can be held in position longer, and therefore more bonding material is used. These wheels are referred to as \u2018hard\u2019 . Sharpness and trueness of the grinding wheel face can be achieved by \u2018dressing\u2019 using an industrial diamond dresser. Dressing with diamonds should always be carried out using a copious supply of coolant which should be turned on before the diamond touches the wheel. The wheel is lowered until it touches the diamond whereupon the diamond is moved across the surface of the wheel, Fig.\u00a010.9. The wheel is then lowered a little and the operation repeated until all the worn grains have been removed, exposing fresh grains and leaving the face flat and true. Vices are used to hold workpieces, but it should be remembered that grinding is usually a finishing operation and so any vice used should be accurate and if possible kept only for use in grinding. Care must also be taken to avoid distortion of the component, as this will be reflected in the finished workpiece. Surfaces required to be ground at right angles can be clamped to the upright surface of an angle plate" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000412_012017-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000412_012017-Figure9-1.png", "caption": "Figure 9. Crack configurations single-edge-crack-at-hole specimen for the calculation of KI (radius, r = 5mm, wide, w = 25mm)", "texts": [ " For this propose, it is needed the weight of unloaded truck of 3000 kg, the vehicle weight during operation is 11000 kg and maximum allowable vehicle load standard is 7500 kg. The load distributed on each wheels is calculated using Eq. (1) [2,18]. IC-DAEM 2018 IOP Conf. Series: Materials Science and Engineering 547 (2019) 012017 IOP Publishing doi:10.1088/1757-899X/547/1/012017 \ud835\udc39 = \ud835\udc5a \u00d7 \ud835\udc54 4 (1) Where m is a vehicle weight of 11000 (kg), g is the gravity of 9, 8 (m/s2). The maximum load leaded by each wheels is 26.950 kN. Figure 9 shows a single-edge-crack-at-hole specimen. Obtaining a value intensity factor (KI) is calculated using Eq. (2) in the basis of data shown in Fig. 9 [19]. \ud835\udc3e\ud835\udc3c = \ud835\udf0e\u221a(\ud835\udf0b\ud835\udc4e)\ud835\udc39\u210e\ud835\udc60 (2) Where \ud835\udc4e is the crack length. The crack length is obtained from the visual observation with the value is 4.9 mm, \ud835\udf0e is the stress that occurs in the spring (specimen). Stress occurring at the spring is calculated using finite element analysis. From finite element analysis as shown in Fig. 10 with a given load of 26950 N, the maximum value of 132.84 (MPa) is obtained. Fhs is a limiting correction factor calculated by referencing to previous research methods [19]. The results of boundary correction factor analysis obtained value of 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001387_cjme.2014.01.211-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001387_cjme.2014.01.211-Figure2-1.png", "caption": "Fig. 2. Mechanical structure of twin ball screw drive", "texts": [ " On the other hand, the dynamic cutting force acts on table, and its vibration will reflect to the machining accuracy. Based on above, the 3D mechanical model to support dynamic analysis of the feed system is proposed. The most common twin ball screw feed drive for precision positioning consists of two ball screws processed as pair, which are driven by two physically separated servomotors through flexible couplings. The ball nut is attached to the table that is constrained by four sliders to move axially on two guides, as Fig. 2 shown. It should be stated that the supporting bearings on the ball screw shaft structure are not expressed in the schematic drawing. The schematic of the mass-spring model considered here is presented in Fig. 3, in which the elements are assumed in the lumped form with 3D mechanical structure parameters. The ball screw shaft structure was endowed with axial stiffness and the guide-slides with support and lateral CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b7213\u00b7 stiffnesses. The parameter definitions of the twin ball screw feed drive system depicted in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003447_s0263574721000229-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003447_s0263574721000229-Figure15-1.png", "caption": "Fig. 15. Measurement configurations of the proposed calibration method. (a) CPA method to identify translational geometric errors. (b) Rotate first three joints to establish E(1)\u2217 = J(1)\u2217 u(1)\u2217 . (c) Rotate last three joints to establish E(2)\u2217 = J(2)\u2217 u(2)\u2217 .", "texts": [ " Then, the pose errors of frame Ftool are measured at these optimized configurations in step 4. Specifically, for the typical calibration method, all joints are involved to form the optimized configurations. Whereas https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574721000229 Downloaded from https://www.cambridge.org/core. Huron University College, on 26 May 2021 at 05:05:45, subject to the Cambridge Core terms of use, available at for the proposed calibration method, first, the translational geometric errors are identified by using the CPA method as shown in Fig. 15(a). Then, as shown in Fig. 15(b), the pose errors of frame Ftool are measured at the optimized measurement configurations of joints 1\u20133, meanwhile, joints 4\u20136 are locked. In this way, the transformation from frame Ftool to F3 is unchanged so that the pose errors of frame Ftool are independent with the unknown parameters of joints 4\u20136. Finally, the poses of Ftool are measured at the optimized measurement configurations of joints 4\u20136 as shown in Fig. 15(c) while joints 1\u20133 are locked. Therefore, frame F3 is held with respect to frame F1 to remove the impact of https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574721000229 Downloaded from https://www.cambridge.org/core. Huron University College, on 26 May 2021 at 05:05:45, subject to the Cambridge Core terms of use, available at unknown parameters of joints 1\u20133 on pose errors of Ftool. During the measurement, the laser beam might be blocked by the robot body and the supporter of SMR at some of the optimized measurement configurations due to the absence of SMR at the back of the supporter (more SMR bases are unavailable in our experiment)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002690_tmag.2020.3017142-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002690_tmag.2020.3017142-Figure1-1.png", "caption": "Fig. 1. Single-sided AFPM machine. (a) Machine configuration. (b) Mechanical Assembly.", "texts": [ " The modal analysis of AFPM machines using different magnetic materials is performed to investigate the machine eigenmodes. Finally, by coupling the results from electromagnetic analysis and modal analysis, the vibration and noise for AFPM machines with different magnetic materials are investigated. The results from vibration and noise analysis verify the established analytical modeling and also shows the influence of different magnetic materials on the machine structural performances. A. Machine Structure Figure 1 shows the configuration and exploded view of a single-sided AFPM machine for the automotive cooling application. This machine is a three-phase, 15-slot/20-pole surface-mounted permanent magnet (PM) machine with a power rating of 2.3 kW and a speed range from 0 to 10,000 rpm. Its outer diameter and axial length are 142 and 32 mm, D Authorized licensed use limited to: Auckland University of Technology. Downloaded on October 03,2020 at 16:00:29 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. > EV-03 < 2 Fig. 2. MMF waveform produced by the tooth-wound stator winding. respectively. The armature winding adopts the tooth-wound approach to reduce the end-winding length. It is not easy to build a machine using ES materials, since the magnetic flux flows perpendicularly into the stator teeth and flows parallel in the stator yoke and rotor yoke, as shown in Fig. 1(a). The rotor yoke and/or the stator yoke can be easily fabricated using laminated ES to increase the effective permeance. For making a stator with the combination of SMC and ES, a hollow cylindrical SMC core firstly is machined to produce the stator teeth and then the SMC stator teeth is jointed with the ES stator yoke [8]. Table I lists the materials used in the AFPM machines for the structural analysis. The SMC and ES materials are used to build the machine stator and rotor core, and the machine housing is made of aluminum alloy", "2 Mode(z, r) = (5,0) Frequency (Hz) 5,511.7 6,257.8 6,368.6, 6,612.8 The eigenmodes of the machine depend on the geometry and materials used for this machine. The natural frequencies fn of these eigenmodes are determined by Young\u2019s modulus E and the mass density \u03c1m of the materials [11]: n mf E (5) Due to the spatial orders here are multiples of 5, the spatial orders above 5 are neglected and detailed model characteristics of these four machines are shown in Table III. The whole machine structure presented in Fig. 1(b) is applied for modal analysis. For a clearer presentation, eigenmodes of the stator are shown in the table. Mode(z, r) means the eigenmode order along the axial and radial directions, and in this paper, the only vibration in the axial direction is considered. It can be found that the natural frequencies of the similar orders of eigenmodes increase due to the use of ES material. Since Young\u2019s modulus of ES material is much larger than that of SMC materials, the AFPM machine using ES materials have higher natural resonant frequencies" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003339_1464420721990049-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003339_1464420721990049-Figure11-1.png", "caption": "Figure 11. Scan trajectory of building cube.17", "texts": [ " Block with dimensions 35 35 35mm3 are fabricated using a printer InssTek MX600 3D equipped with an Ytterbium fiber laser as the heat source with an used power of 417 W at COMTES FHT. It was run with a laser scanning speed of 14.166 mm/s for the filling and contour operations. The nominal layer thickness was 0.25 mm. Powder particles had a size in the range of 20 to 80 mm and were spherical in shape. The laser spot diameter, track overlap, and feed-rate are 0.8 mm, 0.5 mm, and 3 gr/min, respectively. The used scan strategy pattern is shown in Figure 11. In previous investigations, it was observed that the material properties depend on the printing direction and the location of the excised specimens from the surface of the fabricated cube due to different temperature gradients of each layer during solidification process.13,17,18 The schematic of the fabricated block and subsequently dissected samples in order to obtain tensile coupons in multiple orientations are shown in Figure 1. A total number of 15 miniature size tensile test samples were extracted from the block in three different orientations with an electrical discharge machine technique", " Then, the scan parameters from the DED machine (e.g., power, beam locations, beam size diameter, scan speed, and travel distance) were incorporated into the input file through a toolpath\u2013mesh intersection module.20 The module finds intersections between the FE meshes and the scan path from ReplicatorG-0040 software.21 The movement of the heat source was implemented with the use of an event series replicating exactly the movement of the scanning strategy22 and the respective element birth technique follows the laser trajectory as shown in Figure 11. Governing equations. The 3D transient temperature distribution of the cube as well as the build plate can be calculated by solving the transient heat conduction equation, equation1 @ @x k @T @x \u00fe @ dy k @T @y \u00fe @ @z k @T @z \u00feQ \u00bc @ qcpT\u00f0 \u00de @t for T < Tliquidus q L for Tliquidus < T < Tsolidus 8< : (1) where k \u00f0Wm 1K 1\u00de is the thermal conductivity, Q (Wm 3) is the laser power per volume fraction, cp (J kg 1 K 1) is the specific heat, and q (kg=m3) is the density and L (J kg 1) is the latent heat of fusion" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002796_icra40945.2020.9196725-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002796_icra40945.2020.9196725-Figure3-1.png", "caption": "Fig. 3. (a) Cassie is controlled by commanding the listed forces and torques on the pelvis. The pelvis is assumed to be the only component with mass. (b) Cassie can regulate its y position of its pelvis by controlling the desired center of pressure.", "texts": [ " Scalar values are italicized. We treat balancing as positioning and orienting the pelvis without regard for the contact positions and achieve it with Fig. 2. This method assumes the legs are massless, and are used only as a means to produce desired forces. This design results in a controller that is not concerned with maintaining a specific foot placement, and instead administers necessary motor torques to achieve desired pelvis orientations and positions regardless of foot placement as seen in Fig. 3. We use a traditional PD controller to command desired forces to the pelvis. The resulting desired forces are mapped in Cartesian space, and converted to motor torques through multiplication of the Jacobian transpose. We denote the current angles and positions in the world frame as pi = [\u03c6i \u03b8i \u03c8i xi yi zi] T (1) where \u03c6i, \u03b8i, and \u03c8i are angles about the x, y, and z axe, xi, yi, and zi are the Cartesian positions, and subscripts i \u2208 {p r l} are for the pelvis, right foot, and left foot. The desired angles and positions are pid = [\u03c6id \u03b8id \u03c8id xid yid zid] T (2) and are represented in the same order as current angles and positions", " (3), where subscript n \u2208 {P D} denotes proportional and derivative gains. Kn = Kn\u03c6 0 0 0 0 0 0 Kn\u03b8 0 0 0 0 0 0 Kn\u03c8 0 0 0 0 0 0 Knx 0 0 0 0 0 0 Kny 0 0 0 0 0 0 Knz (3) The hand-tuned gains are multiplied by the errors to develop a simple PD controller. The pelvis errors are used to create desired forces for both legs to apply to the ground, so we compute two force vectors. The general equation form is fj = KPpje + KDvje (4) where fj are the torques and forces applied to the pelvis, [Tx Ty Tz Fx 0 Fz] T , from leg j \u2208 {l r} as visualized in Fig. 3a. When a j subscript is present, it applies to only the left or right leg, where as the i subscript can refer to the pelvis, left leg, and right leg. Note, Fy is omitted purposefully with Cassie, where the duty of regulating yp is allocated to a later-described center-of-pressure controller (illustrated in Fig. 3b). In Eq. (5) we develop the individual Jacobian matrix for each leg. BJj = \u2202(Bpj1) \u2202(qj1) . . . \u2202(Bpj1) \u2202(qj5) ... . . . ... \u2202(Bpj6) \u2202(qj1) . . . \u2202(Bpj6) \u2202(qj5) (5) Here, [qi1 . . . qi5] represents motor angles, and [Bpi1 . . .B pi6]T represents the respective foot angles and positions in the pelvis frame (shown in Fig. 4b). The Jacobian transposes are multiplied by pelvis forces, fj , as seen in Eq. (6) to transform individual motor torques. This is done for both the right and left legs independently", " With the few exceptions discussed below, these errors are taken and multiplied by the KP and KD gains to compute pelvis forces. 3) Center-of-pressure Control: The Cassie series robot uses five motors to control the available six Cartesian degrees of freedom per foot. We found directly controlling all Cartesian forces except for the y force to be an effective control scheme. Instead, the y pelvis error is used to compute a desired center of pressure (ycp) between the two feet as illustrated in Fig. 3b and in ycp = yp \u2212Kcpype (9) where Kcp represents a new center-of-pressure gain. Using a simple statics equation we next develop a ratio that determines the necessary weight distribution per leg to achieve the desired ycp, Cr = ycp \u2212 yr\u2211 yj . (10) When multiplied by the weight, this ratio (Cr) not only ensures the weight is always supported by the controller, but also is used to apportion vertical forces to each leg as seen in (11) and (12) Frz = WCr (11) Flz = W \u2212 Frz (12) fjw = [0 0 0 0 0 Fjz] T (13) where W is an estimate of the total weight of the robot. This process is visually described in Fig. 3b. 4) Modifying x Rotation: From the center of the pelvis point to a point internal to the hip motor is a fixed distance, yh, of 0.135 m. Then we determine whether the legs are inside or outside the form factor of the pelvis using Eq. (14). A yf > 0 indicates the feet span is greater than the distance between the left and right hip points. yf = yh \u2212 \u2211 | yj | 2 (14) Eq. (15) uses the yf to determine the qi1 motor angles (reference Fig. 4b) necessary to induce zero roll when the pelvis is centered. These two motors are the only two directly associated with control of the pelvis roll" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure5.8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure5.8-1.png", "caption": "Figure 5.8 Internally threaded ends of length bars", "texts": [ "6(c); (iii) a robust base for converting a gauge-block combination together with the scriber into a height gauge, Fig. 5.6(d); Figure 5.5 Set of gauge-block accessories )c()a( )b()d( Figure 5.6 Assembled gauge blocks and accessories D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 5 Standards, measurement and gauging 5 66 Grade 2 bars are intended as workshop standards for precision measurement of gauges, jigs, workpieces etc. Grades 1 and 2 have internally threaded ends, Fig.\u00a05.8, so that each bar can be used in combination with another by means of a freely fitting connecting screw to make secure lengths. These thread connections are assembled handtight only. The screws also allow the use of a range of accessories. In order to obtain specific lengths, gauge blocks can be wrung to the end faces of the length bar. to assemble any combination. Accessories can be purchased in sets as shown or as individual items. Length bars can be used for greater stability when the use of longer end standards is required" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001985_1.4034768-Figure22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001985_1.4034768-Figure22-1.png", "caption": "Fig. 22 Whirling motions of both defective rotors", "texts": [ "org/about-asme/terms-of-use surfaces. Then pressing force brings about nonuniform stress and bending deformation of rods. Then, counterforce causes the disk\u2019s axis-center to have displacement rb (Fig. 20). This error makes e\u00bc 9.22 lm and rb equals 0.27 lm after pretightening which reaches 3.39 lm at 7500 rpm. Figure 21 shows that the defective rod-fastened rotor bearing system with concentricity error has the similar stability and bifurcation characteristics with previous two kinds of defective rotors. Meanwhile, Fig. 22 shows the whirling orbits of both defective rotors. Figure 23 indicates that the both quasi-periodic motions have similar components: the harmonic frequency (138 Hz and 132.5 Hz), Hopf T frequency (75.6 Hz), and other frequencies. 5.3 Comparison Between Different Defects. The investigated three machining errors have the identical size (10 lm). The dynamic effects are compared in Fig. 24: (1) The flatness error has higher influence on the stability and vibration than the other two machining errors" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure35-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure35-1.png", "caption": "Figure 35. Effect of an umbrella disk", "texts": [ " Thermal gradients and expansions generate thermal stresses in addition to mechanical stress. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. 7.5. Effect Umbrella Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. The temperature rise of a brake disc causes a movement of the friction tracks relative to the initial state. This deformation is called umbrella effect, (Figure 35). The maximum total is deformed at the outer rim of the disc, 284.55 microns it reaches at time t = 3.5 s. An increase in pressure and hence the temperature on a limited contact surface can create this phenomenon (heat deformation), local material fatigue and sometimes cracking of the disk. The umbrella disk effect is not desired because it has a negative influence on the effectiveness of the brakes. Proper operation of a brake under the influence of a heat load is limited by certain events such as thermomechanical: \u2022 Cracking due to the temperature gradient on the slopes of friction, which can cause rupture of the disc, \u2022 The deformation of the disc due to heat (umbrella effect) which influences the contact surface, \u2022 There by reducing the effectiveness of the brake, \u2022 Disc wear and brake pads \u2022 The impact on the environment of the disc (the stirrups, the state oil" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002900_j.tws.2020.107201-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002900_j.tws.2020.107201-Figure6-1.png", "caption": "Fig. 6. Distribution of stress (a) \u03c31 and (b) \u03c32 obtained with Abaqus, and distribution of stress (c) \u03c31 and (d) \u03c32 obtained with SSM.", "texts": [ " As the density of the membrane is 1420 kg/m3 and that of air is 1.29 kg/m3, the added-mass ratio of this membrane is 11.88 computed with Eq. (38). The SSM equations from Section 2 and the free vibration equations from Section 3 are implemented into a program. Besides, the nonlinear finite element analysis is conducted in Abaqus to examine the stress distribution as well. The maximum and the minimum principal stresses are denoted as \u03c31 and \u03c32, respectively. Then the distributions of \u03c31 and \u03c32 are computed with Abaqus and SSM, respectively, which are shown in Fig. 6. The peaks of \u03c31 are located at the four corners of the membrane, and the values of \u03c31 fall steeply inward until form a flat zone in the center, which are depicted in Fig. 6(a) and (c). In Fig. 6(b) and (d), a central peak and four corner peaks of \u03c32 exist in the membrane. However, a tiny zone of \u03c32 is shown negative close to each corner in Fig. 6(b). Besides, a valley zone of \u03c32 exists near each corner as shown in Fig. 6(d). The midpoint of each edge and the center point in the membrane are two feature points, and their stress values are listed in Table 1 for comparison. At both feature points, the values of \u03c31 obtained with SSM are close to those obtained with FEA. Besides, FEA and SSM obtain similar values of \u03c32 at the center point. It is found that a disagreement happens at the midpoint of each edge for \u03c32. In the static analysis, the distribution of \u03c32 obtained with SSM matches the FEA results well in the central zone, but the difference between SSM and FEA results occurs in the edge and corner zones" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002950_tbme.2020.3042115-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002950_tbme.2020.3042115-Figure1-1.png", "caption": "Fig. 1. A) Design of the acoustic stack, and B) Fabricated transducer on the distal end of the steerable guidewire. Scale bar represents 10 mm.", "texts": [ " To achieve reasonable penetration without compromising resolution, a center frequency in the range of 17 MHz was chosen to provide an axial resolution of \u223c92 \u03bcm (17 MHz with 2 cycles). Interventionalists frequently use standard 0.035 in guidewires (0.88 mm outer diameter circular geometry) to image moderately-sized vessels in which chronic occlusions are common (e.g., aortic, carotid, femoropopliteal, or iliac, with diameters in the range of 4\u201310 mm [28]). The 1D Krimholtz-Leedom-Matthaei (KLM) model (PiezoCad, Sonic Concepts, Bothell, WA, USA) [56], [57] was used to design the acoustic stack (Fig. 1A). The single element transducer was composed of an acoustic stack made of highdielectric constant lead zirconate titanate (PZT) (PZT-5H, TRS Technologies, State College, PA, USA) as well as matching and backing layers. Gold was sputtered on each face (200 nm). To improve efficiency of acoustic transmission into and from tissue and increase bandwidth, a matching layer was designed to couple the PZT having high acoustic impedance (34 MRayl) to blood with much lower acoustic impedance (1.5 MRayl). Conductive epoxy (E-Solder 3022, Von Roll Isola Inc", " Radiofrequency (RF) data were filtered using a fourth order Butterworth bandpass filter (10-25 MHz) and processed in MATLAB using a custom script (Mathworks, Natick, MA, USA). Mechanical steering allows for navigation through tortuous vasculature and partially-occluded vessels and also allows acquisition of synthetic aperture data. A steerable robotic guidewire similar to those in [17], [59] was created by micromachining a Nitinol tube of outer diameter 0.88 mm with the Optec femtosecond laser. A Nitinol wire with a diameter of 0.13 mm was used as a tendon and was attached at the distal tip of the tube (Fig. 1B). Applying tension causes bending of the tip of the guidewire in the plane of the notches (setup shown in Fig. 2). The guidewire was held at the proximal end using a 3D-printed fixture (Projet, MJP 5600, 3Dsystems, Rock Hill, SC, USA), and the tendon Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 19,2021 at 06:41:28 UTC from IEEE Xplore. Restrictions apply. was linearly actuated with a lead screw connected to a DC motor (Maxon Precision motors, Taunton, MA, USA) (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002762_j.euromechsol.2020.104125-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002762_j.euromechsol.2020.104125-Figure3-1.png", "caption": "Fig. 3. (a) Configuration during rolling element interaction with outer race defect (Luo et al., 2019a), (b) sketch depicting various angles.", "texts": [ " (2019a), the time separation tsd(or Tsd) was found out analytically using a physics-based approach, while tdi (or Tdi) was found out from recorded signal. Contrary to this, Chen et al.(Chen and Kurfess, 2019) determined both the time separations directly from the signal. In the present study, a modified form of the approach proposed in Luo et al. (2019a) is used. This section is dedicated to the derivation of the relation between time separations and spall length for outer as well as inner race defects. The rolling element center positions during the interaction and geometric relationships are shown in Fig. 3 for outer race defect. The time separation from start to complete destressing in case of outer race defect is given in Eq. (1) (Luo et al., 2019a) tsd = \u03c8b + \u03c8d \u03c9cage (1) where \u03c9cageis cage angular speed, and the angle \u03c8b (Fig. 3(b)) can be obtained from Eq. (2)(Luo et al., 2019a) \u03c8b \u2248 tan(\u03c8b)= bo ( 0.5Dp + 0.5Db ) \u2212 (\u03b4 + 0.5cl) (2) where bo is semi-width of contact area in the tangential direction (Fig. 2 (d and e) and Fig. 3(a)), Dpis pitch diameter of the bearing,Db is rolling element diameter, clis radial clearance (Fig. 3(b)), \u03b4is total deformation of rolling element-raceway contact in the radial direction. The semiminor dimension bo of contact area between the outer race and rolling element at position \u03d5 is given as (Harris, 1996) bo = b* o [ 3F\u03d5 2 \u2211 \u03c1o ( 1 \u2212 \u03be2 I EI + 1 \u2212 \u03be2 II EII )]1 / 3 (3) where b* ois dimensionless parameter dependant on the curvature of bodies in contact, F\u03d5is load acting on the rolling element at angular position \u03d5 (Fig. 2 (c) and (d)), \u2211 \u03c1o is the sum of curvature of outer racerolling element contact, \u03be is Poisson\u2019s ratio, E is the modulus of elasticity, while the subscripts I and II indicate two bodies in contact" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000436_acc.2019.8815013-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000436_acc.2019.8815013-Figure8-1.png", "caption": "Fig. 8. CFD Simulation of Wind Disturbance", "texts": [ " This figure also demonstrates that the tilt servos are abetting the TRQ system to reach the desired states. The second case involves providing wind disturbances to the TRQ system and studying the response of the system to the proposed sliding mode controller inputs. To generate real world wind disturbance scenarios, CFD simulations were first run using ANSYS Fluent 18.1 on a quarter CAD model (designed using SOLIDWORKS 2015) in the in-house Super Computing facility utilizing 16 cores and 128GB RAM. One such simulation result for a wind speed of 10m/s is shown in Figure 8. These simulations were used to generate drag forces acting on the TRQ system for certain wind velocities. The wind velocity to drag force co-relation simulation results are provided in Table IV. The rationale behind running the simulations on a quarter model and not a complete model is that we are interested in the moments rather than just the forces. To demonstrate the robustness of the sliding mode controller, the wind force (Fw in the Z direction is Fwc\u03b8) is applied only to the motor 1 which relates to a random pitching force acting on the system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001537_j.triboint.2015.12.035-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001537_j.triboint.2015.12.035-Figure3-1.png", "caption": "Fig. 3. Schematic of the bearing.", "texts": [ " [24] show with a time consuming sophisticated model including thermal effects on top foil and bumps that a maximum of 10% difference in temperature exists between the upper surface of base block and top foil. 8. Viscosity variations Temperature variations have a direct impact on lubricant viscosity. For gas, we used an explicit formulation which gives the molecular viscosity as a function of the temperature [19]: ~\u03bc \u00bc ~T 1:5 \u2219 1\u00fe Su T0 ~T \u00fe Su T0 \u00f032\u00de The Sutherland number Su is a constant which depends on the lubricant composition, which value can be find in the literature. 9. Bearing geometry A sketch of the bearing in consideration is given in Fig. 3 (see Appendix A for details). The dimensionless film thickness is: ~h \u00bc 1\u00fe\u03f5i cos\u03b8 \u00f033\u00de 9.1. Foil deflection: Heshmat model The variation of the film thickness, \u210e, is due to the eccentricity ratio \u03b5i and the deflection of the foil ~\u03c9t under the imposed hydrodynamic pressures developed between the bearing clearance and it is given by: ~h \u00bc 1\u00fe\u03f5i cos\u03b8\u00fe\u03c9t \u00f034\u00de where ~\u03c9t is the dimensionless elastic deformation of the foil structure under the imposed hydrodynamic pressure. This deformation ~\u03c9t [20] depends on the bump dimensionless compliance ~\u03b1t and the average pressure across the bearing width (Fig", " The FDM theory is based on simple principles but it can solve rather complex TEHD problems in an efficient and accurate way. Therefore, a lot of recent studies use FDM to solve hydrodynamic or TEHD problems [21,22]. Besides, there is no doubt that FDM is very convenient when working with simple geometries such as the plain journal bearing or GFB simplified profiles. The global numerical algorithm is outlined in the flowchart represented in Fig. B1. 13. Results and discussion: 3D THD analysis We use a GFB which characteristics and running conditions are described in Fig. 3 and Table 1. 13.1. Pressure fields The pressure field (Figs. 5\u20138) is clearly depending on the axial location and high pressure area is centered on the mid-length location in this case at second lobe whatever the used running conditions. In the high pressure area, in almost one fourth of the zone at the bearing edges, the pressure increase is at least 20% smaller than at mid-length. We took for the calculation 180 points in circumferential direction, and 31 points in axial direction. Fig. 5(a and b) depicts the steady-state pressure fields calculated for a loaded journal bearing operating at 15 N, shaft speed 120,000 and 180,000 R" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001589_0954410015612499-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001589_0954410015612499-Figure1-1.png", "caption": "Figure 1. Missile\u2013target engagement geometry.", "texts": [ " The set _V \u00bc 0 consists of the S\u00f0t\u00de, e1\u00f0t\u00de\u00f0 \u00de \u00bc \u00f00, 0\u00de, and the only invariant set inside S\u00f0t\u00de, e1\u00f0t\u00de\u00f0 \u00de \u00bc \u00f00, 0\u00de is the origin S\u00f0t\u00de, e1\u00f0t\u00de, e2\u00f0t\u00de\u00f0 \u00de \u00bc \u00f00, 0, 0\u00de. Thus, the asymptotic convergence of S\u00f0t\u00de, e1\u00f0t\u00de and e2\u00f0t\u00de to zero is assured and the stability of closed-loop system (26) will be guaranteed. Guidance law design In this section, first the model of guidance loop dynamical system is introduced and then SMG law to intercept target is designed. Modeling of guidance loop In this section, a model for guidance loop is formulated. Consider a two-dimensional interceptor and target engagement as shown in Figure 1. It is assumed that the missile and the target are point masses moving in plane. Then, the missile\u2013target engagement model shown in Figure 1 can be described by the following nonlinear differential equations11 _R \u00bc VT cos\u00f0q T\u00de VM cos\u00f0q M\u00de \u00f030\u00de _q \u00bc 1 R VT sin\u00f0q T\u00de \u00fe VM sin\u00f0q M\u00de\u00f0 \u00de \u00f031\u00de _ M \u00bc 1 VM AM \u00f032\u00de _ T \u00bc 1 VT AT \u00f033\u00de where R is the relative range between target and interceptor q is the LOS angle with respect to a reference axis, _q is the line of sight rate, R _q is the relative lateral velocity, VM and VT represent the interceptor and target velocities respectively, M and T represent the Fight path angles of the target and missile, and AM and AT represent the interceptor and target accelerations, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001507_iccas.2015.7364708-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001507_iccas.2015.7364708-Figure1-1.png", "caption": "Fig. 1: The components of the harmonic drive CSF-20", "texts": [ " We have been developing the harmonic reduction gear, called KCSF-20, by tak ing the reverse engineering of CSF-20 model (Harmonic Drive Systems, Inc.) and its dynamic characteristics was evaluated through the modal testing. The Campbell dia gram was used to analyse the vibrational characteristics of CSF-20 according to angular velocity of the shaft. 2. HARMONIC REDUCTION GEAR 2.1 Configuration Harmonic reduction gear is in general consist of three basic components which are flex spline, circular spline and the wave generator. Fig. 1 shows the components of the CSF-20 model. The wave generator includes an el liptical shaped-bearing plug and a ball bearing, and func tions as an input driver. When the bearing plug rotates around the circular spline, the elastic deformation of the flex spline occurs and the reduction gear obtains the gear reduction ratio by engaging the additional two teeth. Hence, it is essential to analyse dynamic characteristics of the flex spline in order to validate the reduction mechanism. 2.2 Modal testing Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure3.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure3.6-1.png", "caption": "Figure 3.6 Surface table", "texts": [ " Marking out polar co-ordinates requires not only accuracy of the dimension along the radial line but accuracy of the angle itself. As the polar distance increases, any slight angular error will effectively increase the inaccuracy of the final position. The possibility of error is less with rectangular co-ordinates, and the polar co-ordinate dimensions shown in Fig. 3.1 could be redrawn as rectangular co-ordinates as shown in Fig. 3.5. In order to establish a datum from which all measurements are made a reference surface is required. This reference surface takes the form of a large flat surface called a surface table (Fig. 3.6) upon which the measuring equipment is used. Surface plates (Fig. 3.7) are smaller reference surfaces and are placed on a bench for use with smaller workpieces. For general use, both surface tables and surface plates are made from cast iron machined to various grades of accuracy. For Datum Figure 3.3 Datum centre lines Datum Figure 3.4 Datum edge and centre line D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 3 3 Marking out When the workpiece has to be positioned at 90\u00b0 to the reference surface, it can be clamped to an angle plate (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure10.11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure10.11-1.png", "caption": "Figure 10.11 Wheel-collet assembly", "texts": [ "1, the diamond is lowered in contact with the top of the wheel and is traversed across the wheel surface using the handwheel (7). An attachment of this kind saves time through not having to position and remove the diamond for each dressing nor having to lower the wheel for dressing and to raise it again to continue grinding. It is impossible to produce good-quality work on any grinding machine if the wheel is out of balance, thus setting up vibrations through the spindle. The wheel is mounted on a collet, the complete assembly being removable from the spindle, Fig.\u00a010.11. The wheel spigot upon which the wheel is located has a taper bore to accurately locate on the spindle nose. The wheel flange locates on the wheel spigot and is held by three screws which, when tightened, securely hold the wheel between the two surfaces. On the outer D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 10 Surface grinding 10 151 Full width guards are fitted to the front of the machine table to protect the operator in the event of a workpiece flying off a magnetic chuck and from coolant spray" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.13-1.png", "caption": "Fig. 3.13 Euler sequence: Left\u2014The xyz (fixed) system is shown in blue, and the XYZ (rotated) system is shown in red. The line of nodes, labeled N, is shown in green (from Wikipedia). Right The corresponding gimbal", "texts": [ "27) When R is given, the Helmholtz angles (\u03b8H , \u03c6H , \u03c8H ) can be using \u03b8H = arcsin(Ryx ) \u03c6H = \u2212 arcsin ( Rzx cos \u03b8H ) \u03c8H = \u2212 arcsin ( Ryz cos \u03b8H ) . (3.28) Yet another sequence to describe 3-D orientation is common in theoretical physics and mechanics and in other technical literature, and often referred to as Euler sequence.5 In order to describe the movement of a spinning top rotating on a table, or of the earth during its rotation around the sun (see Fig. 3.12), three angles are needed: the intrinsic rotation (\u03b3 ), nutation (\u03b2), and precession (\u03b1). Using these three angles, the orientation of the spinning object is described by (see Fig. 3.13) \u2022 a rotation about the z-axis, by an angle \u03b1, \u2022 followed by a rotation about the rotated x-axis, by an angle \u03b2, and \u2022 followed by a rotation about the twice-rotated z-axis, by an angle \u03b3 . REuler = Rz(\u03b1) \u00b7 Rx (\u03b2) \u00b7 Rz(\u03b3 ). (3.29) This leads to the parametrization REuler = \u23a1 \u23a3 \u2212 sin \u03b1E cos\u03b2E sin \u03b3E + cos\u03b1E cos \u03b3E \u2212 sin \u03b1E cos\u03b2E cos \u03b3E \u2212 cos\u03b1E sin \u03b3E sin \u03b1E sin \u03b2E sin \u03b1E cos \u03b3E + cos\u03b1E cos\u03b2E sin \u03b3E \u2212 sin \u03b1E sin \u03b3E + cos\u03b1E cos\u03b2E cos \u03b3E \u2212 cos\u03b1E sin \u03b2E sin \u03b2E sin \u03b3E sin \u03b2E cos \u03b3E cos\u03b2E \u23a4 \u23a6 (3.30) 5The expression Euler angles should be used very carefully: sometimes, these angles represent the Euler sequence, but often that expression is also applied when the nautical sequence is actually used" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002534_012146-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002534_012146-Figure3-1.png", "caption": "Figure 3. Three-dimensional sketch of the Rotor", "texts": [ " Unlike previously considered designs [10], the prototype of 500 kW generators has additional reserve copper windings of the armature 6 and excitation 16. In the event of a malfunction of the cryogenic supply system, the electric generator can remain operational at 10 percent of power. Copper windings are also necessary in the first stage for testing of the prototype. Figure 2 shows that the windings of the armature right packages are rotated by 15 geometric degrees. Three-dimensional sketch of the rotor is shown in figure 3. Rotor packages are made of sheets of electrical steel; permanent magnets are in their closed grooves. The insignificant amount of pole scattering fluxes is compensated by a decrease in the air gap. The hollow shaft has a lower mass and sufficient rigidity to increase the critical speed. A variant of the elaborated principal design is shown in figure 4. 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001533_0142331215619972-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001533_0142331215619972-Figure4-1.png", "caption": "Figure 4. Lift vector components during turn.", "texts": [ " The main tasks of the guidance law are to: generate a smooth and bounded fref , using measurements of the states y, x and f, to keep the lateral error small in the presence of disturbances; regulate track error to zero by graceful bank-to-turn maneuvers, without excessive overshoot and maintain the general flight direction WP1! WP2; keep the guidance output magnitude jfref j below the maximum value fmax; maintain stability during flight; fulfill performance and robustness requirements. Aerial vehicles use a component of the aerodynamic lift to generate lateral accelerations to correct lateral errors. Lateral acceleration is normally produced by rolling or banking the vehicle so that the lift vector is tilted in the direction of the required turn (the bank-to-turn maneuver), this is depicted in Figure 4 . Lift is divided into two components, one balances the centrifugal force and the other balances the weight of the vehicle L cosf=mg, L sinf= mV 2 R \u00f03\u00de where m is the mass, g is the gravitational acceleration, V is the ground speed, and R is the radius of turn of the vehicle. For a coordinated turn we have tanf= V 2 Rg \u00f04\u00de During a steady turn V =R _x, so equation (4) takes the form tanf= V _x g \u00f05\u00de Now since _x E = _x _x R , therefore we have tanf= V ( _x E + _x R ) g \u00f06\u00de or _x E = g V tanf _x R \u00f07\u00de where _x R is input from the mission plan" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000270_978-3-030-24741-6_10-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000270_978-3-030-24741-6_10-Figure2-1.png", "caption": "Fig. 2. Compliant foil model demonstrator. (A) Technical sketch of simplified bioinspired geometry (carnivorous plant trap morphology). (B) The foil was folded at the connection of the triangles (lobes) to the rectangle. An acrylic microscope slide was attached to the rectangle as (modifiable) backbone. (C) Left: Closure movement of the lobes in response to an applied downward force to the \u201cears\u201d. Right: Snap-buckling of the backbone in response to an applied bending force to the backbone, resulted in an opening of the lobes.", "texts": [ " Additionally, we will show the possibility of energy storage and movement locking via hydrogel based actuators. The novel actuator systems presented in this study can serve as a basic outline for developments of smart bioinspired and autonomous demonstrators with tailored motion sequences and movement speeds within Cluster of Excellence livMatS. The basic morphology of the two carnivorous plants (two lobes connected by a midrib) was abstracted into a simple planar geometry consisting of two triangular lobes vertically connected by a rectangle that was joined by two circles as \u201cears\u201d for actuation (Fig. 2A). The sketch was printed onto a foil and cut out (coated copying film 5001476 with 0.1 mm thickness from Streit GmbH & Co). The foil was buckled at the connection of the triangles to the rectangle. An acrylic plastic microscope slide was fixed under the rectangle (same size) as a rigid backbone, to enhance structural mechanical stability (Fig. 2B). When the \u201cears\u201d were pressed down manually, the lobes underwent a closing movement due to the curvature of the \u201cears\u201d and their attachment points to the lobes (kinematic coupling, as present in the Aldrovanda trap). Alternatively, when the \u201cears\u201d were fixed and the rigid backbone was moved upwards, this will result in the same movement. Applying an additional force exerted from below towards the center of the backbone will result in a bending of the backbone and cause a snap buckling motion (as present in Dionaea), which re-opened the lobes (Fig. 2C). For the actuation of the closing movement three different actuator systems were used. The demonstrator was actuated via pneumatic cushions (pneumatic model) (Fig. 3A\u2013D), a shape memory alloy (SMA) spring was attached to the \u201cears\u201d (thermal model) (Fig. 3E\u2013G), or magnets attached to one \u201cear\u201d and actuated contactless by a magnetic stirrer (magnetic model) (Fig. 3H\u2013I). The pneumatic model was based on three pneumatic cushions in a frame (Fig. 3A) to which the \u201cears\u201d of the demonstrator are fixed at the edge (adhesive tape)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001170_0278364914551773-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001170_0278364914551773-Figure5-1.png", "caption": "Fig. 5. Rocky7 design. The rocker\u2013bogie passive suspension system is intended to better support motion over uneven terrain. All wheels are driven and the front two are also steered. Figure inspired by Tarokh and McDermott (2005).", "texts": [ "comDownloaded from reproduces the articulated wheel equation expressed in world coordinates: vw c = vw v \u00bd w rv c \u00d7 vw v \u00bd w rs c \u00d7 w vv s \u00f092\u00de Therefore, it is possible to use a systematic recursive approach in coordinatized form as well, if one is willing to use the coordinatized form of the transport theorem to perform the differentiations. When it comes to computer implementation, coordinates are necessary anyway, so the two approaches are equivalent implementations of different notations for the same process. Rocky7 is probably the most complicated WMR kinematics problem that has been addressed successfully in the literature. Here, rather than attempt a closed-form solution, we will formulate a recursive numerical solution for the position and velocity kinematics. The overall design of the rover is illustrated in Figure 5. The main rocker angle r is antisymmetric, having opposite values on each side of the body on uneven terrain (r1 = 2r2). The two smaller rockers on either side are passive (and independent) at their \u2018\u2018bogie\u2019\u2019 joints b1 and b2. This setup provides exactly the three passive degrees of freedom needed to allow the six wheels to remain in contact with the terrain. Coordinate frames are assigned according to the DH convention as in (Tarokh and McDermott, 2005) in Figure 6. Only the left side is shown" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002388_s00170-020-05218-9-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002388_s00170-020-05218-9-Figure2-1.png", "caption": "Fig. 2 Geometry of overlap between two successive layers in multi-layer single-track deposition: a ideal or theoretical case; b actual or practical case", "texts": [ " Furthermore, no study has been made to investigate the dimensional and geometrical accuracy of the components manufactured by the \u03bc-PTAAM process. Therefore, the present paper presents development and experimental validation of a theoretical model for evaluation of total height and width of multi-layer single-track deposition by the \u03bc-PTAAM process and its use to experimentally investigate tolerances, and dimensional and geometrical deviations of metallic components manufactured by the \u03bc-PTAAMprocess. Figure 2 a shows the schematic of an ideal or theoretical overlap SOR between two successive deposition layers, whereas Fig. 2 b depicts schematic of actual or practical deposition in which overlap between two successive layers is not feasible resulting an increase in area of the second deposition layer to S\u2019T\u2019P\u2019Q\u2019R\u2019. The following assumptions were made to develop the theoretical model for prediction of total height and width of multi-layer single-track deposition by the \u03bc-PTAAM process. & Geometry of profile of single-track deposition in the first layer is assumed to be semi-elliptical as per findings of Jhavar et al. [27] and from the second layer onward it has been assumed to be circular as per assumption of Ding et al. [29] with its center being at the top of the previously deposition layer, i.e., the point \u201cO\u201d as shown in Fig. 2 a. & Each combination of parameters of the \u03bc-PTAAMprocess yields uniform deposition profile throughout its length. The profile of a deposited track is symmetrical about the vertical axis. & The overlap SOR between the first and second deposition layers is equal to the non-overlap STPQR as height and width of an individual layer are in range of 1\u20133.5 mm and areas \u201cTOS\u201d and \u201cQOR\u201d in non-overlap \u201cSTPQR\u201d (Fig. 2a) are significantly smaller than the cross-sectional area of multi-layer single-track deposition. Overlap SOR and non-overlap STPQR are assumed as a semi-circle of radius \u201cr\u201d for theoretical or ideal case of the deposition. But during the actual deposition process, the second layer is deposited over the first layer without any overlap as it is physically infeasible. Therefore, the overlap is added to non-overlap, resulting in increase in radius of deposition from \u201cr\u201d to \u201cr\u2032\u201d and enlargement of the profile to \u201cS \u2032 T \u2032 P \u2032Q \u2032 R\u2032\u201d of the second layer of deposition as depicted in Fig. 2 b. Width \u201cw\u201d (mm) and height \u201ch\u201d (mm) of single-layer single-track deposition have been evaluated using Eq. 1 and Eq. 2 given by the thermal model of Nikam et al. [30] which has been developed using the principles of energy balance and heat transfer and considering thermal properties of the substrate and deposition materials, and three parameters of the \u03bc-PTAAM process namely input power to the micro-plasma \u201cP\u201d (which is equal to product of DC voltage \u201cV\u201d (volts) and current \u201cI\u201d (amperes) supplied to the micro-plasma); travel speed of micro-plasma torch \u201cVt\u201d (mm/s); and volumetric material deposition rate \u201cVd\u201d (mm3/s)", " w \u00bc 2 \u03b7P\u2212Vd\u03c1dCpd Tmd Ti\u00f0 \u00de 17:8\u03c1sC * ps Tms\u2212Ti\u00f0 \u00de ffiffiffiffiffiffiffiffiffi \u03b1sV t p erfc 1\u00f0 \u00de 1 exp ffiffi p p \u00fe 1ffiffi p p 2 4 3 5 2=3 \u00f01\u00de h \u00bc 4 1\u2212D\u00f0 \u00deVd \u03c0 Vt w \u00f02\u00de Where, \u03b7 is the thermal efficiency of the micro-plasma transferred arc (%); \u03c1d is the density of the deposition material (kg/mm3); Cpd is the specific heat of the deposition material (J/kg K); Tmd is the melting temperature of the deposition material (K); Ti is the ambient temperature (K); \u03c1s is the density of the substrate material (kg/mm3); Cps* is the modified specific heat of the substrate material (J/kg K); Tms is the melting temperature of substrate material (K); \u03b1s is the thermal diffusivity of the substrate material (mm2/s); and D is dilution (%). Values of thermal efficiency for micro-plasma arc \u201c\u03b7\u201d as 60% and dilution \u201cD\u201d as 15% as per assumption of Sawant et al. [31]. Equating the area of a semi-elliptical crosssection of the first deposition layer with the area of a circular cross-section of the second deposition layer in an ideal case of deposition is shown in Fig. 2 a: \u03c0 2 h w 2 \u00bc \u03c0 r2\u21d2r \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 0:25 h w p \u00f03\u00de As per the assumption, consideration of geometry of the actual multi-layer single-track deposition requires that the area of the overlapping SOR between the first and second deposition layers be added to the area of non-overlapping STPQR between them which will increase the radius of the 2nd deposition layer from \u201cr\u201d to \u201cr\u2032\u201d and enlarge its profile to S \u2032 T \u2032 P \u2032 Q \u2032 R\u2032 as depicted in Fig. 2 b. \u27f92 \u03c0 r2 2 \u00bc \u03c0 2 r 02 \u21d2r 0 \u00bc ffiffiffiffiffiffi 2 r p \u27f9r 0 \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffi 0:5 h w p \u00f04\u00de Width of multi\u2212layer single\u2212track deposition0W 0 \u00bc 2r 0 \u00bc ffiffiffiffiffiffiffiffiffiffiffi 2 h w p \u00f05\u00de Total height (H) of multi-layer single-track deposition for (n) numbers of deposition layers can be estimated from following relation: H \u00bc h\u00fe n\u22121\u00f0 \u00der0 \u00bc h\u00fe n\u22121\u00f0 \u00de ffiffiffiffiffiffiffiffiffiffiffiffiffiffi 0:5 h w p \u00f06\u00de Previously optimized values of the \u03bc-PTAAM process parameters (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002262_b978-0-12-815020-7.00005-9-Figure5.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002262_b978-0-12-815020-7.00005-9-Figure5.6-1.png", "caption": "FIGURE 5.6 Workpiece design for slot end-milling experiments. (A) 3D view of workpiece with machined eight slots and eight drilled thermocouple holes, and (B) top and side views of workpiece with dimensions and hatched section.", "texts": [ " Level Factors Cooling methods Speed (RPM) Feed (IPM) [mm] 0 Emulsion 1000 6 [150] 1 MQL 1500 12 [300] 2 LN2 2000 18 [450] block, Block B, has the experimental runs (slots) for LN2 cooling strategy (experimental runs 4, 8, and 9) on one half of the block and on the second half of the block three slots were machined for a comparative evaluation of MQL, LN2, and combined (MQL + LN2) cooling strategies using the identified optimum machining parameters from the ANOVA of fractional factorial experimental design. The workpiece geometry and dimensions are shown in Fig. 5.6. Four-flutes uncoated solid carbide helical end-mills of 0.5 in. (12.7 mm) diameter, 0.5 in. (12.7 mm) shank diameter, 1 in. (25.4 mm) flute length, and 3 in. (76.2 mm) overall length with 0.03 in. (0.76 mm) corner radius were used. A corner radius of 0.03 in. was chosen to give minimum cutting force values as reported by Okafor and Aramalla [13]. New endmills were used for each experimental run (slot) to exclude tool wear effect, and a total of 12 end-mills were used for all experiments. The spindle speed range and feed rate range were selected to suitably cover the recommended range of machining conditions for titanium alloy, while cooling methods and cryogenic temperature of \u221215 \u00b0C were selected from recommendations in the published literature", " The cutting force components were acquired at every pass when the end-mill has traveled about 1 in. (25.4 mm) into the slot. A new end-mill was used per slot and forces acquired at pass numbers 1 and 2, and 7 and 8 were treated as replicates for purpose of ANOVA and error estimation for sharp and worn tools, respectively. Forces acquired for pass numbers 1, 2, 4, 5, 7, and 8 were used for indirect tool wear monitoring per machined slot and cooling strategy. The design of the workpiece block and slots is shown in Fig. 5.6. Eight 0.125 in. (3.175 mm) diameter, 1 in (25 mm) deep holes were drilled on the front sides of all the titanium Ti-6Al-4V blocks at the middle of each slot at 0.25 in. (6.35 mm) from the bottom of the block. An ungrounded K-type thermocouple probe of 0.125 in. (3.125 mm) diameter was inserted into the drilled holes to ensure snug fit. The thermocouple probe was connected to a National Instrument NI USB-9211A Data Acquisition Device (DAQ) for thermocouple and the DAQ device was connected to a desktop computer" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002854_j.engfailanal.2020.105028-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002854_j.engfailanal.2020.105028-Figure13-1.png", "caption": "Fig. 13. Structural optimization.", "texts": [ " It is usually to increase the cross-sectional area of the object to improve the fatigue resistance of the object, or to change the material properties by common design methods. Therefore, the priority is to increase the cross-section area to improve the fatigue safety factor of the weak area of the connecting rod by this design scheme. At the same time, two oil passages in the connecting rod are changed into one to improve the fatigue strength. In order to ensure the same lubrication conditions, the diameter of the oil passage is increased by 1.4 times. The optimized structure is shown in Fig. 13. It can be seen that the cross-sectional area and transition area of the connecting rod Z. Pan and Y. Zhang Engineering Failure Analysis xxx (xxxx) xxx have been modified, the modified cross-sectional area is increased by about 30%. The fatigue safety factor of key points in the new scheme is analyzed by applying the same load and boundary constraints. From Fig. 14, the fatigue safety factors of points 1 and 2 are greatly improved. After increasing the cross-sectional area of the connecting rod, the fatigue safety factor of the connecting rod is improved" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000613_j.mechmachtheory.2019.103718-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000613_j.mechmachtheory.2019.103718-Figure5-1.png", "caption": "Fig. 5. PTPM. (a) Computer-aided design (CAD) model and (b) schematic diagram.", "texts": [ " U ij is the strain energy stored in the joint, where the subscripts i and j denote that the joint is the j th component in the i th limb. W ij can be expressed using W i based on the force analysis: W i j = Q i j W i , (25) where Q ij denotes the mapping matrix from W i to W ij . Substitution of Eq. (25) into (24) gives U i j = 1 2 W T i Q T i j \u02dc C i j Q i j W i = 1 2 W T i C i j W i , (26) and C i j = Q T i j \u02dc C i j Q i j , (27) where C ij is the compliance matrix of the j th component in the i th limb. A PTPM, shown in Fig. 5 , consists of three identical limbs that connect a moving platform to a fixed base. Each limb consists of a lower arm and an upper arm. Each lower arm represents a translational actuator and each upper arm consists of a planar four-bar parallelogram shape. The three limbs constrain the moving platform output to produce translational motion. The origin of the fixed coordinate frame O - XYZ is located at the center of the base, the Z axis lies perpendicular to the base plane, the Y axis points in the direction of OD 1 , and the X axis is determined by the right-hand rule", " 6 and the equilibrium equation can be expressed as W = J f , (28) where J represents the overall constraint wrenches that can be expressed as J = [ J 11 , J 12 , J 21 , J 22 , J 31 , J 32 ] T and J ij = [ l i , pA ij \u00d7 l i ], while l i and pA ij represent the unit vectors of the side shafts A i 1 B i 1 and pA ij , respectively. f = [ f 11 , f 12 , f 21 , f 22 , f 31 , f 32 ] T represents the magnitudes of the overall constraint matrix. It is important to note here that the PTPM is a non-overconstrained PM and the magnitudes of the overall constraint wrenches can thus be obtained from the equilibrium equation directly without the deformation compatibility equation. f = J \u22121 W = D W , (29) where D = J \u22121 . When e i 1 is considered to be flexible and all other elastic components are rigid, as shown in Fig. 5 (b), the total strain energy of the PTPM is equal to the strain energy that is stored in elastic component e i 1 . U = U i 1 = 1 2 f T i 1 c s f i 1 = 1 2 ( B 2 i \u22121 W ) T c s B 2 i \u22121 W = 1 2 W T B T 2 i \u22121 c s B 2 i \u22121 W , (30) where B 2 i \u22121 denotes row 2 i \u22121 of matrix B . The elastic deflection of the moving platform that is caused by elastic component e i 1 can be obtained using Castigliano\u2019s second theorem. i 1 = \u2202U \u2202W = B T 2 i \u22121 c s B 2 i \u22121 W = C\u0304 i 1 W , (31) where c s is the flexibility coefficient of the spherical joint that can be determined using ANSYS commercial software" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001554_0959651815595909-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001554_0959651815595909-Figure2-1.png", "caption": "Figure 2. Two-degree-of-freedom model of vehicle dynamic.", "texts": [ "sagepub.comDownloaded from In this section, lateral model of the vehicle is presented. In Figure 1, the modeling of vehicle dynamics performed by Rajamani and Doumiati is illustrated.4,18 The authors have demonstrated that the bicycle model with 2 degrees of freedom (DOFs) consists of some essential characteristics of a vehicle lateral dynamics.18,19 The 2-DOF model is represented by the vehicle lateral position y and the vehicle yaw angle c. An approximation of lateral vehicle model is shown in Figure 2. In the construction of this model, two left and right front wheels of car model and two left and right rear wheels have been, respectively, depicted with single wheel in point A and single wheel in point B. Point C shows gravity center of the vehicle.19b is an angle that the vehicle velocity vector has to the longitudinal axis in gravity center and is called the sideslip angle of the vehicle. Lateral motion equation for vehicle dynamics is taken from Newton\u2019s second law, which is the initiator of the main part of dynamics analysis and is as follows4,19 may =Fyf +Fyr \u00fe Fdy \u00f01\u00de where Fdy is the lateral force disturbance, m is the total mass of the vehicle, and ay = dy2=dt2 is the vehicle inertial acceleration at gravity center, which consists of two components ay = \u20acy vx _c \u00f02\u00de The first component \u20acy is an acceleration produced via vehicle motion along with the y-axis, and the second component is called centripetal acceleration" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002258_s00170-020-04953-3-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002258_s00170-020-04953-3-Figure11-1.png", "caption": "Fig. 11 The novel CNC hourglass worm grinding machine. a Machine structure; b Grinding machine", "texts": [ " 10, \u03c6bc is the step length of match angle \u03c6c, \u03b5 is the accuracy of solution, and the step-by-step searchmethod is used to search the optimal match angle. In order to illustrate the validity and practicability of the virtual rotation center grinding method and the measuring principle, an hourglass worm grinding machine is developed, and a planar enveloping hourglass worm is ground, as well as the measure test and performance test are carried out. According to the virtual rotation center grinding principle of planar enveloping hourglass worm tooth surface, an hourglass worm grinding machine is produced [26\u201328] as shown in Fig. 11. There are seven kinematics axes in this hourglass worm grinding machine. The grinding movements of this hourglass worm grinding machine are the X axis, Z axis, B axis, and C axis, and these four axes\u2019motions are full closed loop control. The adjustment movement of this hourglass worm grinding machine are the A axis and the Y axis. The A axis is the adjustment motion of the grinding wheel angle, while the Y axis is the height adjustment motion of the grinding wheel. The geometry parameters of the planar enveloping hourglass worm is shown in Table 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001755_978-3-319-26327-4-Figure6.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001755_978-3-319-26327-4-Figure6.2-1.png", "caption": "Fig. 6.2 Omnidirectional camera model using Geyer and Daniilidis\u2019 representation (Geyer and Daniilidis 2000)", "texts": [ " However, the knowledge of these functions as well as their gradient are necessary for the actual computation of the equations in (6.7). Herein, we will briefly overview two models that we used for experimental evaluation of the proposed method: the first one is the classical catadioptric camera model of Geyer and Daniilidis (2000) and the second one is the model of Scaramuzza (2006a) who derived a general polynomial form of the internal projection valid for any type of omnidirectional camera. Let us first see the relationship between a 3D point X and its projection x in the omnidirectional image I (see Fig. 6.2). Note that only the half sphere on the image plane side is actually used, as the other half is not visible from image points. The camera coordinate system is in S, the origin (which is also the center of the sphere) is the effective projection center of the camera and the z axis is the optical axis of the camera which intersects the image plane in the principal point. To represent the nonlinear (but symmetric) distortion of central catadioptric cameras, (Geyer and Daniilidis 2000) projects a 3D point X from the camera coordinate system to a virtual projection plane P through the virtual projection center CP = (0, 0,\u2212\u03be)T as xP = h(X) = \u23a1 \u23a2 \u23a3 X1 X2 X3 + \u03be \u221a X2 1 + X2 2 + X2 3 \u23a4 \u23a5 \u23a6 The virtual plane P is then transformed in the image plane I (see Fig. 6.2) through the homography HC as x = HC xP HC = KC RMC , where KC includes the perspective camera parameters (taking the picture of the mirror), R is the rotation between camera and mirror, while MC includes the mirror parameters\u2014see (Geyer and Daniilidis 2000) for details. Herein, we will assume an ideal setting: no rotation (i.e. R = I) and a simple pinhole camera with focal length f and principal point (x0, y0) yielding HC = \u23a1 \u23a2 \u23a2 \u23a3 f \u03b7 0 x0 0 f \u03b7 y0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a2 \u23a3 \u03b3 0 x0 0 \u03b3 y0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a6 where \u03b3 = f \u03b7 is the generalized focal length of the camera-mirror system and \u03b7 is the mirror parameter" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001140_speedam.2014.6871911-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001140_speedam.2014.6871911-Figure7-1.png", "caption": "Fig. 7. Arbitrary Injection based estimation scheme block diagram.", "texts": [ " From (23), (24) and (29) the only remaining machine parameter for position estimation is the mean admittance Y\u03a3. The resulting current slope for an applied voltage vector represents the admittance for that voltage excitation. From [22] the admittance for a certain switching cycle can be calculated as Y\u0302tX = \u2206isHF T vss vss Tvss , X = 1, 2, 3. (30) Since the injected voltages are symmetrically distributed the mean admittance estimation equals Y\u0302\u03a3 = Y\u0302t1 + Y\u0302t2 + Y\u0302t3 3 . (31) The complete AI based estimator is shown in Fig. 7 in block diagram format. The new hybrid estimation scheme derived herein combines the AF and AI methods using a hysteresis approach shown in Fig. 8. The regions were chosen using the following criteria: \u2022 At least 500 r/min mechanical rotor speed to enter AF (this allows fast enough PLL synchronisation) \u2022 At least 200 r/min to maintain AF estimation \u2022 At least 1 A d-axis saturation current (is = 2 A at \u03b3 = 60\u00b0) for AF \u2022 A 500 mA current buffer for hysteresis Switching between estimation schemes only occurs when crossing the edges of the hysteresis region in the direction of the arrows shown" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000105_s10958-019-04227-8-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000105_s10958-019-04227-8-Figure1-1.png", "caption": "Fig. 1", "texts": [ " The coefficients of the polynomial are determined from the initial and final conditions of operation, while the coefficients of the trigonometric series are determined from the solution of the corresponding problem of nonlinear programming. We perform numerical experiments and analyze, on their basis, the influence of the configurations of the manipulator and the parameters of the trigonometric series on the characteristics of the suboptimal process. The proposed investigation is a subsequent development of the procedure of parametric optimization in the problems of optimal control over nonlinear mechanical systems [5, 7\u201313, 20, 25]. Consider a mechanical model of a two-link manipulator schematically shown in Fig. 1. The manipulator consists of two solids (links) OA and AB connected by a cylindrical hinge A . The link OA is connected with the immobile base with the help of a cylindrical hinge O and the grip with a load is placed at the end of the second link. The axes of the hinges O and A are directed along the vertical. The hinges are regarded as perfect and the gripper (with the load) is modeled by a point mass m concentrated at the point B . Under the action of the moments of controlling forces u1 and u2 applied to the hinges O and A , respectively, the manipulator undergoes plane-parallel motions in the horizontal plane OXY " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003429_j.tws.2021.107585-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003429_j.tws.2021.107585-Figure4-1.png", "caption": "Fig. 4. Boundary conditions for (a) open apex shell, (b) closed apex shell.", "texts": [ " (38) are not suitable for the shells with variable radii of curvature. The modified trigonometric series is designed to describe normal displacements characterized by the disturbances in the area of the shell edges that diminish relatively quickly along the shell meridian. Aside from those disturbances, the displacements remain constant which is true for the shells with constant radii of curvature. The boundary conditions applied on the displacement functions vary, depending on whether the closed apex or open apex shell is considered (Fig. 4). In the case of closed apex i.e. \ud835\udf111 = 0 following boundary conditions are assumed: ( ) ( ) \ud835\udc51\ud835\udc63 \ud835\udf111 = 0, \ud835\udc51\u210e \ud835\udf112 = 0, \ud835\udf171(\ud835\udf111) = 0. (40) The boundary conditions provided in Eqs. (40)\u2013(42) consider separated shell structures. To achieve structural compatibility in the junctions of deformed pressure vessel additional boundary conditions have to be applied. The compatibility equations i.e. boundary conditions for the junction of the ellipsoidal and cylindrical shell (Fig. 5) have the form: \ud835\udc51\ud835\udc63 ( \ud835\udf11\ud835\udc522 ) = \ud835\udc51\ud835\udc63 ( \ud835\udf09\ud835\udc502 ) , \ud835\udf171 ( \ud835\udf11\ud835\udc522 ) = \ud835\udf171 ( \ud835\udf09\ud835\udc502 ) " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003850_j.jallcom.2021.161608-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003850_j.jallcom.2021.161608-Figure1-1.png", "caption": "Fig. 1. (a) Schematic illustration for preparation and assembly of Ni(OH)F@GF, (b) Schematic illustration of the flexible microelectrode sensor. (c) Photographs of flexible microelectrode sensors.", "texts": [ " Cyclic voltammograms (CVs) and Electrochemical impedance spectroscopy (EIS) of the fiber-shaped sensors were measured on an electrochemical workstation (CHI 760E, CH Instrument) with a threeelectrode system in 0.1 M NaOH solution. An Ag/AgCl electrode and a platinum wire were used as the reference electrode and the counter electrode, respectively. The Ni(OH)F@GF composite fiber was used as the working electrode. The preparation process of Ni(OH)F@GF core-sheath fiber and the fabrication process of flexible fiber-shaped glucose sensor are presented in Fig. 1. As shown in Fig. 1a, The Ni2+, F- and OH- released from Ni(CH3COOH)2, NH4F and CO(NH2)2, respectively, can react with each other to form nanorods on the surface of graphene fiber (GF). Usually, Ni(CH3COO)2 and CO(NH2)2 can react with each other to form nickel hydroxide nanorods during the hydrothermal process. However, in the present of NH4F, the F- from NH4F also participates in the reaction, forming a kind of fluorine doping nickel hydroxide. the resulting nanorod material was denoted as Ni(OH)F, and the composite fiber of graphene decorated with Ni(OH)F nanorods was signed as Ni(OH)F@GF. The synthesis of Ni(OH)F can be attributed to the following equations [32]. ++Ni(CH COO) Ni 2[CH COCO]3 2 2 3 (1) + + ++NH F H O NH H O H F4 2 3 2 (2) + + + +CO(NH ) 3H O CO 2OH 2[NH ]2 2 2 2 4 (3) + ++Ni F OH Ni(OH)F2 (4) The flexible microelectrode sensor was fabricated by sealing the Ni(OH)F@GF composite fiber in a transparent PVDF capillary tube. As shown in Fig. 1b, one end of Ni(OH)F@GF composite fiber was connected to a copper wire with conductive silver paste and placed in a PVDF tube, and both ends of the PVDF tube are sealed with epoxy resin. Fig. 1c are photographs of the flexible microelectrode sensor with Ni(OH)F@GF, and the exposed part of the Ni(OH)F@GF composite fiber at one end of the microelectrode is about 5 mm. Compared glass capillary tube microelectrodes [33], our PVDF tube microelectrode shows better toughness and flexibility. Fig. 2a shows the SEM image of the pristine graphene fiber. The diameter of the pristine graphene fiber is about 80 \u00b5m. The highresolution SEM images of pristine graphene fiber (Fig. 2b,c) show graphene wrinkles which were formed during the preparation process of graphene fiber" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000488_chicc.2019.8866201-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000488_chicc.2019.8866201-Figure2-1.png", "caption": "Fig. 2: AGV Kinematics model", "texts": [ " To avoid collisions, consider the size and shape of the AGV, assuming that the width of the AGV is w, simplifying the road model for the feasible area, and the feasible area of the AGV becomes the area between f \u2032 1(x) = f1(x) \u2212 w/2 and f \u2032 2(x) = f2(x) + w/2 .Suppose fb(x) is the lateral position of the XOY rear axis, \u03c8 is the yaw angle, and the distance from the AGV rear axis to the front axis is l. The lateral position of the front end of the simplified AGV is fa(x) = fb(x) + l sin\u03c8 . The feasible area of the AGV is: f \u2032 2(x)\u2212 l sin\u03c8 \u2264 fa(x) \u2264 f \u2032 1(x)\u2212 l sin\u03c8 (1) f \u2032 2(x) \u2264 fb(x) \u2264 f \u2032 1(x) (2) The AGV kinematics model is shown in Fig. 2. (Xr, Yr) and (Xf , Yf ) represent the coordinates of the AGV rear axis and the front axis, respectively, l is the distance from the AGV rear axis to the front axis, and \u03c8 is the AGV yaw angle. \u03b4f is the front wheel deflection angle of the AGV, and vr and vf respectively represent the AGV rear axle center speed and the front axle center speed. As can be seen from Figure 2, the axis speed of the rear axle is: vr = X\u0307r cos\u03c8 + Y\u0307r sin\u03c8 (3) The AGV steering process model diagram is shown in Fig. 3. R is the AGV rear wheel steering radius, P is the instantaneous rotational center of the AGV, and C and D are the AGV rear axle axis and the front axle axis, respectively. It is assumed that the AGV clockwise turning radius is the same as the road curvature radius during the steering process. The kinematic equations of the front and rear axes are: X\u0307f sin(\u03c8 + \u03b4f )\u2212 Y\u0307f cos(\u03c8 + \u03b4f ) = 0 (4) X\u0307r sin\u03c8 \u2212 Y\u0307r cos\u03c8 = 0 (5) According to the front and rear wheel geometry, we can know: Xf = Xr + l cos\u03c8 (6) Yf = Yr + l sin\u03c8 (7) Combined with the above formula, the kinematics model of AGV is: X\u0307r = vr cos\u03c8 (8) Y\u0307r = vr sin\u03c8 (9) \u03c8\u0307 = vr tan \u03b4f/l (10) By linearizing it, the discrete linear time-varying model of AGV kinematics can be obtained as follows: x\u0303(k + 1) = Ax\u0303(k) +Bu\u0303(k) (11) among them, x\u0303 = \u23a1 \u23a3 x\u2212 xr y \u2212 yr \u03c8 \u2212 \u03c8r \u23a4 \u23a6A = \u23a1 \u23a3 1 0 \u2212vr sin\u03c8rT 0 1 vr cos\u03c8rT 0 0 1 \u23a4 \u23a6 (12) u\u0303 = [ v \u2212 vr \u03b4f \u2212 \u03b4f,r ] B = \u23a1 \u23a3 cos\u03c8rT 0 sin\u03c8rT 0 tan \u03b4f,rT l vrT l cos2(\u03b4f,r) \u23a4 \u23a6 (13) Where x\u0303 is the deviation of the current trajectory from the desired trajectory, u\u0303 is the deviation of the control variable, k is the sampling time, and T is the sampling period" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003178_j.jsv.2015.06.037-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003178_j.jsv.2015.06.037-Figure9-1.png", "caption": "Fig. 9. The mode shapes of healthy rotating impeller\u2013shaft-bearing assembly: (a) the shaft-dominated mode and (b) the impeller-dominated mode.", "texts": [ " 8(a). An interest veering phenomenon appears at high frequency range between the two frequency whirl lines due to the crack, such as the results shown in Fig. 8(b). In Fig. 8, both the eigenvalues of healthy and cracked assembly are included. Fig. 8(a) shows the low frequency region of the Campbell diagram. It can be seen that the 60 mm crack leads to negligible changes in system's lower eigenvalues. This is because the low-order mode shapes of the assembly are mostly shaft-dominated as shown in Fig. 9(a). Local stiffness reduction due to crack may not significantly affect system's effective stiffness. At the high frequency region, the differences between eigenvalues of the healthy and cracked assembly become quite obvious as shown in Fig. 8(b). The relative difference between the eigenvalues may reach up to 20 percent. The high-order mode shapes are impeller-dominated, such as the mode shape shown in Fig. 9(b). A local crack can lead to remarkable reduction in impeller's effective stiffness, as well as the assembly's natural frequencies. Moreover, the high-order eigenvalues of cracked assembly exhibit complex trend with rotating speed. At the range of low rotation speed, such as 0\u20139000 rev/min, the larger natural frequency of the cracked assembly is nearly equal to the lower frequency of its healthy counterpart, whereas the lower frequency of cracked assembly and the lower frequency of healthy assembly have similar trend at the range of high speed", " The orbit shown in Fig. 13(a) is relatively simple and contains only one nodal point. This may be owing to that some lower-order shaft-dominated bending modes are excited by the EO1 excitation. As is known in turbomachinery that a certain engines order of forces only excite modes with the same number of nodal diameters (NDs) for cyclic symmetric structures. For the assembly, there exist numerous modes that are coupled modes between the 1ND modes of impeller and the bending of shaft, such as the mode shape shown in Fig. 9(b). In contrast, the orbit shown in Fig. 13(b) is much more complex than that shown in (a). The three obvious nodal points illustrate the intricate deflection of shaft. As the blade passing frequency tends to be very large, high-order modes will be excited. Then, the effects of crack on the transverse response of nodes on shaft axis at the speed of 9000 rev/min are presented in Fig. 14. The response in each figure contains three cases, which are referred to as \u201cHealthy\u201d, \u201cLinear\u201d and \u201cNonlinear\u201d" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001083_1.4027926-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001083_1.4027926-Figure7-1.png", "caption": "Fig. 7 Faults in 6214 ball bearing of compressor rotor experimental rig", "texts": [], "surrounding_texts": [ "4.1 Faults Produced on Ball Bearings. To study the casing responses caused by ball bearing faults, in this study, faults are produced on ball bearings using wire-electrode cutting technology. These faults are produced on the inner ring, outer ring, and ball of the compressor experimental rig\u2019s 6214 ball bearing and the aero-engine experimental rig\u2019s 6205 ball bearing, which are shown in Figs.7 and 8. The ball bearing dimensions are listed in Table 1. The ball bearing faults\u2019 characteristics frequencies can be computed as follows:\n(1) Outer race\nfoc \u00bc Z 2 1 d D cos a fR (1)\n(2) Inner race\nfic \u00bc Z 2 1\u00fe d D cos a fR (2)\n(3) Rolling ball\nfbc \u00bc Z\n2\nD d 1 d D cos a\n2 \" #\nfR (3)\n(4) Cage\nfc \u00bc 1 2 1 d D cos a fR\n(4)\n4.2 Faults Experiments of Ball Bearings. The bearing faults experiments are carried out. The test-site photo of the compression rotor rig is shown in Fig. 9, and the measurement points\u2019 explanation is listed in Table 2; the test-site photo of the aeroengine rotor rig is shown in Fig. 10, and the measurement points\u2019 explanation is listed in Table 3.\nVibration signals are collected by means of the USB9234 data acquisition card of the NI Company, the 4805 type ICP acceleration sensors of B&K Company are used to pick up the acceleration signals, and the eddy current sensors are used to measure the rotating speeds. The sampling frequency is 10.24 kHz.\n4.3 Wavelet Envelope Analysis for Ball Bearing Fault\n4.3.1 Basic Principles of Wavelet Envelope Analysis. In this study, a signal analysis to determine ball bearing faults is carried out by means of a wavelet envelope spectrum analysis. Chen [2] provides a reference for the detailed process of this algorithm. The essence of ball bearing fault diagnosis based on the wavelet packet is to take advantage of its bandpass filter characteristics, and to decompose the signals using appropriate wavelet functions so as to obtain an appropriate resonance frequency band. Then, by means of envelope demodulation, low frequency envelope signals that only contain the fault characteristic information are obtained. Its spectrum is the wavelet envelope spectrum, in which the fault characteristic frequencies of the ball bearings can be found out.\n4.3.2 Wavelet Envelope Spectrum Analysis of Ball Bearing Fault Signals\n4.3.2.1 Experiment Analysis Based on Compressor Rotor Experimental Rig\n(1) Feature extraction for inner ring faults\nFigures 11\u201314, respectively denote the time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope spectra of the bearing house\u2019s response and the casing\u2019s response to inner ring faults in the 6124 ball bearing. The experimental rotating speed is 1793 rpm (29.88 Hz). The number of balls is 10, the other ball bearing\u2019s parameters are listed in Table 1, and the inner ring characteristic frequency can be calculated by formula (2) as fic\u00bc 5.8974 29.88\u00bc 176.24 Hz. From Fig. 11, the vibration acceleration response value of the casing is larger than that of the bearing housing, which shows that the impulse vibration of the bearing fault is effectively transmitted to the casing. This result is in full agreement with the experimental result for the impulse response. From Fig. 12, it can be found out that there are many resonance peaks in the frequency spectrum, and the signals are very weak in the low frequency segments, the signal\u2019s frequency spectrum from 0 Hz to 500 Hz are shown in Fig. 13, from which, the characteristic frequency of the inner ring fault and its modulation frequency fr can be\nJournal of Dynamic Systems, Measurement, and Control NOVEMBER 2014, Vol. 136 / 061009-3\nDownloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "basically seen. The wavelet envelope spectrum is shown in Fig. 14, by comparison with Fig. 11, the characteristic frequency of the inner ring fault and the modulation frequency fr are more distinctly shown in Fig. 14. Obviously, the inner ring fault features can be all extracted from the acceleration signals of the bearing house and the casing by means of the wavelet envelope spectrum analysis and frequency spectrum. (2) Feature extraction of outer ring faults The time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope\nspectrum of the bearing house responses and casing responses of the 6124 ball bearing outer ring faults are respectively shown in Figs. 15\u201318. The experimental rotating speed is 1826 rpm (30.43 Hz). The number of balls is 11, the other ball bearing\u2019s parameters are listed in Table 1, and the outer ring characteristic frequency can be calculated by formula (1) as foc\u00bc 4.5879 30.43\u00bc 140 Hz. From Fig. 15, the vibration acceleration response value of the casing is larger than that of the bearing housing, which shows that the impulse vibration of the bearing fault is effectively transmitted to the casing. This result is in full agreement with the experimental result for the impulse response. From Fig. 16, it can be found out that there are many resonance peaks in the frequency spectrum, and the\nTable 1 Ball bearing dimensions (in mm)\nType Diameter of inner ring Diameter of outer ring Thick Diameter of ball Pitch diameter\n061009-4 / Vol. 136, NOVEMBER 2014 Transactions of the ASME\nDownloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "signal\u2019s spectrum from 0 Hz to 500 Hz is shown in Fig. 17, from which, the characteristic frequency of the outer ring fault can be distinctly seen. The wavelet envelope spectrum is shown in Fig. 18, by comparison with Fig. 17, the characteristic frequency of the outer ring fault are also distinctly shown in Fig. 18, obviously, outer ring fault features can be all effectively extracted from the acceleration signals of the bearing house and casing by means of frequency spectrum and the wavelet envelope spectrum analysis. In addition, no modulation frequency components appear in the frequency spectrum and the wavelet envelope spectrum. (3) Feature extraction for ball faults The time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope spectrum of the bearing house\u2019s response and casing\u2019s response for the 6124 ball bearing\u2019s ball faults are respectively shown in Figs. 19\u201322. The experimental rotating speed is 1827 rpm (30.43 Hz). The number of balls is 11, the other ball bearing\u2019s parameters are listed in Table 1, the ball fault characteristic frequency can be calculated by Formula (3) as fbc\u00bc 5.3919 30.43\u00bc 164.1 Hz, and the cage rotating frequency can be calculated by formula (4) as fc\u00bc 0.4103 30.43\u00bc 12.5 Hz. From Fig. 19, the vibration acceleration response value of the casing is larger than that of the bearing housing, which\nFig. 10 Test-site photo of aero-engine rotor experimental rig\nTable 3 Explanation of the measurement points of the aeroengine rotor experimental rig\nChannel Measurement variable Measurement point position MOD3, CH1 Rotating speed Coupling MOD3, CH2 Acceleration Bearing house MOD4, CH1 Acceleration Casing top MOD4, CH2 Acceleration Casing right MOD4, CH3 Acceleration Casing bottom MOD4, CH4 Acceleration Casing left\nJournal of Dynamic Systems, Measurement, and Control NOVEMBER 2014, Vol. 136 / 061009-5\nDownloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_22_0000919_chicc.2015.7260287-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000919_chicc.2015.7260287-Figure2-1.png", "caption": "Fig. 2: Structure of the x-z inverted pendulum", "texts": [ " In Section 3, the simulation results of the x-z inverted pendulum are given in different conditions. Section 4 gives the conclusions of this paper. The customary inverted pendulum on a cart driven by the horizontal control force is shown in Fig.1. In Fig.1, the control action is based on the horizontal displacements of the cart. Compared with customary inverted pendulum, the xz inverted pendulum which is mounted on a cart driven by the combination of the horizontal and vertical control forces is introduced in Fig.2. In Fig.2, the inverted pendulum is mounted on a little platform that can move up and down through a vertical force and can move with the cart in the horizontal direction. We can conclude the following remarks though comparison of two types of inverted pendulum. (1) Pure vertical control force can not realize stabilization of the inverted pendulum[14,15]. Combination of the horizontal and vertical control forces make the inverted pendulum have more flexibilities for control design. (2) The x-z inverted pendulum is more like the control of practical situations than the customary inverted pendulum, such as stabilizing a pole on one finger or practical control of the rocket" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001842_oceansap.2016.7485341-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001842_oceansap.2016.7485341-Figure1-1.png", "caption": "Fig. 1. The scheme location of hoists", "texts": [ " The successful use of DUV under the conditions of sea disturbance requires taking some measures to reduce its impact on a DUV. This problem can be solved by means of an automatic control system of a descent underwater vehicle. To control a DUV under the conditions of sea disturbance we suggest using a system with a boat hoist (BH) (for lowering-lifting of a descent underwater vehicle) and shockabsorbing hoist (SAH) (for oscillation damping caused by sea disturbance). The SAH is located on a DUV (Figure 1) follow. The speed of vertical longitudinal movement of a DUV being in lowering-lifting mode in this system should be equal to the cable traverse speed at the output block of a boat hoist. It is suggested ensuring the attitude stability mode of a DUV under the conditions of sea disturbance on the basis of the cable tension deviation compensation measured near a DUV. To describe the control system of a DUV mathematically some specific features of the link \u00abcable-DUV\u00bb should be taken into consideration" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001020_s00170-015-7033-2-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001020_s00170-015-7033-2-Figure8-1.png", "caption": "Fig. 8 Grinding of the first flank angle", "texts": [ " According to the requirements of the dip angle of grinding wheel, flank angle, and flank, the grinding process of the flank is as follows: 1. Determine the initial position of flank grinding according to the mathematical model of cutting edge. 2. Rotate the workpiece around a-axis and position the workpiece along X and Z directions. 3. The grinding wheel feeds along z-axis and approaches to the grinding position. 4. The grinding wheel rotates around y-axis by \u03b11 (the first flank angle) and feeds down along y-axis to make the grinding surface closed to the ground workpiece, as shown in Fig. 8. 5. The workpiece rotates around a-axis, and simultaneously the grinding wheel feeds along X direction which is a Then the grinding wheel performs a similar process to form the second flank angle, as shown in Fig. 9. If more flank angle or auxiliary groove is to be added, the above-mentioned steps should be repeated. Through the grinding simulation of the double-circular-arc torus milling cutter, the fabrication process of the milling cutter is completed, as shown in Fig. 10. The cutting edge\u2019s shape and grinding trajectory is verified" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001689_0040517514547210-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001689_0040517514547210-Figure4-1.png", "caption": "Figure 4. A geometric model of the final triangle.", "texts": [ " Here, w and h are the width and height of the primary triangle, is the apex angle of the primary triangle, /2 is the inclination angle of the yarn spinning tension, i is the angle between the ith fiber and the central fiber, and n is the fiber number in each fiber strand, that is, there are 2n fibers in the Siro-spinning triangle. O1 and O2 are the middle point of the nip line of each primary triangle, C1 and C2 are the twisting point of each fiber strand, and F1 and F2 are the spinning tension of each fiber strand. The final spinning triangle is shown in Figure 4. Here, C is the twisting point of the final spun yarn, F is the yarn spinning tension, which can be measured using a tension tester directly, and is the apex angle of the final triangle. According to force balance in Figure 4, we know F1 cos 2 \u00fe F2 cos 2 \u00bc F \u00f01\u00de F1 sin 2 \u00bc F2 sin 2 \u00f02\u00de Then, we have at BROWN UNIVERSITY on May 20, 2015trj.sagepub.comDownloaded from F1 \u00bc F2 \u00bc F sin 2 sin \u00f03\u00de In order to investigate the mechanical performance of a Siro-spinning triangle by using the FEM, the following assumptions are made.1\u20133 Assumption 1.1,3 The cross-section of all fibers is a circle with identical diameters. All fibers are gripped between the front roller nip and the twisting point. The velocity of fibers in the spinning triangle is constant, and the delivery velocities of fibers and yarn are the same" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001435_s10439-014-1185-3-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001435_s10439-014-1185-3-Figure2-1.png", "caption": "FIGURE 2. In silico model for determining, via optimization, the lumped rotational stiffnesses (K1 and K2) and damping coefficients (B) for the elbow and shoulder (Left: sagittal plane, Right: transverse plane). The black dots denote spherical joints at wrist (extension), elbow (flexion), shoulder (extensor and adduction) and the sternoclavicular joint to ground.", "texts": [ " A 3-D, sagittally-symmetric, fourlink (including hand, forearm, upper arm and clavicle), lumped parameter, musculoskeletal representation was employed using equations of the form: si \u00bc Kihi\u00fe Bi _hi \u00fe Ii\u20achi, where si denotes the applied torque of each joint i, hi is the angular displacement of each joint i, _hi is the angular velocity of each joint i, \u20achi is the angular acceleration of each joint i, Ii is the calculated moment of inertia of each segment i, Bi is the rotational damping coefficient, and Ki is the rotational stiffness coefficient for each joint i. The moment of inertia was calculated by measuring the length of the limb and knowing the location of the center of mass of forearm and upper arm. The subject-specific anthropometric data are described in the next paragraph. The model arm segments were assumed to be connected by four frictionless spherical joints at wrist, elbow, shoulder and sternoclavicular joints (Fig. 2). Segment anthropometric, mass, and inertial properties were scaled to each subject\u2019s height and weight based upon the literature.34 The model arm muscles were represented by a bilinear torsional spring and linear damper placed in parallel at the elbow, and again at the shoulder and sternoclavicular joints. The bilinear behavior for each joint was characterized using two rotational stiffnesses (K1 and K2) and a single damping coefficient (B) in an optimization algorithm. Joint kinematics, including wrist angle, elbow angle, shoulder angle and wrist displacement, were measured as the arm was end loaded in a proximal direction in each impulsive loading trial" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000096_042019-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000096_042019-Figure4-1.png", "caption": "Figure. 4. The design scheme with a rotor-sided thrust bearing CC:1 - impeller; 2 - rotor; 3 - radial bearings; 4- thrust bearing;5- thrust disk", "texts": [ " MMBVPA IOP Conf. Series: Journal of Physics: Conf. Series 1158 (2019) 042019 IOP Publishing doi:10.1088/1742-6596/1158/4/042019 intercushionchannel skew. Cross section: a - along the rotational axis and thrust disk; b - A-A; c - B-B MMBVPA IOP Conf. Series: Journal of Physics: Conf. Series 1158 (2019) 042019 IOP Publishing doi:10.1088/1742-6596/1158/4/042019 Dynamic processes in lubricating boundary layers and the bearing are determined by the behavior of the rotor, disposed in radial and thrust bearings (Fig. 4). Assuming independence of the axial movement from the radial, one can describe the behavior of the rotor in the axial direction. For this purpose, on the basis of the second law of Newton write equation of rotor dynamics in the axial direction: 2 . . , 2 d \u0443d disp m P Fr d (1) where mr \u2013 rotor mass is constant and is determined during construction of the compressor; . .\u0443d disp - coordinate disc displacement along the rotor axis; - time; \ud835\udc43 \u2212carrier bearing ability; F - foreign unsteady gas dynamic force of the CC, acting on the rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000959_s00502-014-0272-3-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000959_s00502-014-0272-3-Figure4-1.png", "caption": "Fig. 4. Structure of an eight-pole double-sided axial flux machine with internal stator and a three-phase tooth-coil winding with q = 1 2 slots per pole and phase [3]: (a) side view, (b) stator front view, (c) rotor front view with surface-mounted trapezoidal permanent magnets", "texts": [ " It has two stator cores, which carry the stator winding and one internal rotor disc. The rotor disc consists of permanent magnets, which are arranged as a ring shape. The magnets are carried by a nonmagnetic and non-conductive carrier construction to avoid eddycurrents due to the stator magnetic field. Due to the two stators, no rotor iron yoke is required to guide the flux. The flux passes directly through the PM, which reduces the axial machine length between the two air-gaps. 26 heft 1.2015 \u00a9 Springer Verlag Wien e&i elektrotechnik und informationstechnik Figure 4 shows the double-sided AFM with internal stator (AFIS). The AFIS-machine consists of two rotor discs, which surround the internal stator core. The stator carries the tooth coil winding, which gives a short radial outer and inner winding overhang. In this type no stator yoke is needed as the stator teeth guide the flux via the two air-gaps directly to the two rotor sides [8]. Hence also a short axial length is possible. Three machines were electromagnetically designed with a 3D FiniteElement-Program: one RFM with an outer rotor, one AFM with an internal stator and one AFM with an internal rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000433_acc.2019.8815065-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000433_acc.2019.8815065-Figure1-1.png", "caption": "Fig. 1. Insertion process by 6-DOF manipulator.", "texts": [ " Fei proposed error recovery strategies for connectors mating based on force sensor [23]. These approaches give less weight to precision because positional error can be cancelled by deformable components. So, they are not suitable for our targeting precise assembly. The purpose of this research is to develop a new insertion method that enables a versatile robot system to execute assembly task with the clearance smaller than 10\u03bcm. The concrete target task presented in this paper is that a robotic hand holds a deformable thin ring and put it around a shaft (see Fig. 1). This task appears frequently in manufacturing scenes including fabrication of cylindrical components, such as a piston and a piston ring in reciprocating engine, a rotor and a shaft bushes of flat motor or hard dDisk drive (HDD), a lens and a flange of zoom lens system, and so on. The inside dimension of the ring and the outside dimension of the shaft are \u03c650H8(50+0.039 0.000 ) and \u03c650g7(50\u22120.009 \u22120.034), respectively. Therefore, the narrowest clearance is 9\u03bcm in our target assembly task. In order to accomplish the assembly target, we propose a passive alignment principle (PAP) for reduction of positioning errors of robot and a new mating and insertion algorithm using the PAP" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003746_s10409-021-01089-9-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003746_s10409-021-01089-9-Figure2-1.png", "caption": "Fig. 2 B\u00e9zier volume and its control points", "texts": [ " B represents the Bernstein basis function, which has the following definition: (5) S h 1 = (2 + 1)( \u2212 1)2(2 + 1)(2 + 1)\u22154, S h 2 = l ( \u2212 1)2(2 + 1)(2 + 1)\u22154, S h 3 = \u2212(2 + 1)( \u2212 1)2(2 \u2212 1)(2 + 1)\u22154, S h 4 = \u2212l ( \u2212 1)2(2 \u2212 1)(2 + 1)\u22154, S h 5 = (2 + 1)( \u2212 1)2(2 \u2212 1)(2 \u2212 1)\u22154, S h 6 = l ( \u2212 1)2(2 \u2212 1)(2 \u2212 1)\u22154, S h 7 = \u2212(2 + 1)( \u2212 1)2(2 + 1)(2 \u2212 1)\u22154, S h 8 = \u2212l ( \u2212 1)2(2 + 1)(2 \u2212 1)\u22154, S h 9 = \u2212 2(2 \u2212 31)(2 + 1)(2 + 1)\u22154, S h 10 = l 2( \u2212 1)(2 + 1)(2 + 1)\u22154, S h 11 = 2(2 \u2212 3)(2 \u2212 1)(2 + 1)\u22154, S h 12 = \u2212l 2( \u2212 1)(2 \u2212 1)(2 + 1)\u22154, S h 13 = \u2212 2(2 \u2212 3)(2 \u2212 1)(2 \u2212 1)\u22154, S h 14 = l 2( \u2212 1)(2 \u2212 1)(2 \u2212 1)\u22154, S h 15 = 2(2 \u2212 3)(2 + 1)(2 \u2212 1)\u22154, S h 16 = \u2212l 2( \u2212 1)(2 \u2212 1)(2 \u2212 1)\u22154, \u23ab \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23ac \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23ad (6)M = \u222bV STSdV . (7)r(u, v,w) = 3\u2211 i=0 1\u2211 j=0 1\u2211 k=0 Bi,3(u)Bj,1(v)Bk,1(w)Pi,j,k, Figure\u00a02 shows a B\u00e9zier volume and its control points. In order to transform the geometry model into a finite element mesh that is capable for dynamic simulation, the transformation matrix that can convert the B\u00e9zier volume control points into the SBE48 nodal coordinates is required. Let P represent the vector form of the control points: According to the property of the B\u00e9zier volume, the position and gradient vectors of the first 4 nodes in SBE48 can be formulated as following: where l is the element length" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003639_j.jfoodeng.2021.110707-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003639_j.jfoodeng.2021.110707-Figure2-1.png", "caption": "Fig. 2. Proposed apparatus design: (a) the overall CAD model, (b) the CAD model of the robotic hand, and (c) the actual developed prototype.", "texts": [ " For measuring the friction coefficient, the robotic hand was made to perform a grasping and lifting motion, while the block bar was placed on the top of the target to prevent it from being lifted. Consequently, this induces a slippage between the robotic hand and the target, and the friction coefficient was determined using the tangential and normal forces measured by the 6-axis force sensor. By controlling the applied grasping force, its relationship with the friction coefficient can be investigated. The developed apparatus is shown in Fig. 2. Autodesk Fusion 360 (educational license, Autodesk Inc., CA) was used to construct the CAD models of the apparatus. The apparatus consists of a robotic hand for grasping, two linear stages for translating the robotic hand and the block bar in the vertical direction, a triangle-shaped block bar, and aluminum frames (Fig. 2a). The robotic hand was constructed using a DC motor (DCX26LGBKL24V, Maxon, Switzerland), a rack and pinion mechanism, two fingers, and a 6-axis force sensor (CFS018CA101U, Leptrino, Japan) placed under the fingertip. Furthermore, the fingertip has a slide slot, which allows for easy change of the fingertip surface. The robotic hand can realize point-to-point (PTP) control and continuous path (CP) control based on the feedback from the motor encoder (ENX10EASY1024IMP, Maxon, Switzerland). The PTP control is used to adjust the initial grasping position of the fingertip, and the CP control is utilized to generate the desired indentation path for measuring viscoelasticity", " Two linear stages driven by stepper motors (23HS6430, MotionKing, China) were used to vertically translate the robotic hand and the block bar, and the largest translation speed can be achieved at 50 mm/s. The block bar blocks the motion of the food target while it is being lifted; this induces slippage between the target food and the robotic fingertip, which helps measure the friction coefficient. Aluminum frames were used to assemble the components to construct the apparatus. The prototype is depicted in Fig. 2c. In this study, we employed the standard linear solid model of the Maxwell form (Lin, 2020) to simulate the viscoelasticity of tempura products. This model was chosen because it has the least number of parameters which can be easily estimated through one-time grasp-and-hold test. Nonlinear behavior of tempuras is not considered because we usually do not tend to generate large deformation on the tempuras during handling tasks to avoid damaging the food products. In addition, we did not consider the plastic deformation that will eventually remain in tempuras because it is difficult to capture both force and residual Z" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001911_speedam.2016.7525881-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001911_speedam.2016.7525881-Figure1-1.png", "caption": "Fig. 1. An axial Halbach rotor. Fig. 2. The 2-D cylinderical surface region for the 2-D", "texts": [ " The equations are then solved using an integral approach. By following the above two steps it is shown in this paper that highly accurate 3-D field and torque equations for a magnetic coupling can be determined. The resulting torque equation is easy to compute as it only needs one integral to be evaluated. Although the approach presented here is for a Halbach AMC the presented approach could be extended to model AMCs that contain magnetic steel [7, 8, 12, 13]. II. 2-D ANALYTICAL MODEL The model of an axial Halbach rotor is shown in Fig. 1. It has an inner radius, ri, an outer radius, ro and an axial length dI. The rotor magnets are magnetized along the axial direction as shown. The field created by the Halbach rotor is first solved at radius rc. It is assumed that at radius rc the radial field is 978-1-5090-2067-6/16/$31.00 \u00a92016 IEEE zero (due to symmetry). Therefore at r=rc the magnetic field is assumed to be fully described by ( , , ) 0 ( , , ) ( , , )c c z cr z B r z B r z B r z (2) Therefore the field B will exist only on a 2-D cylindrical surface as illustrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002475_s40430-020-02295-5-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002475_s40430-020-02295-5-Figure10-1.png", "caption": "Fig. 10 Distribution of sound field in circumferential direction under different speeds", "texts": [ "\u00a03, the 24 field points were selected to analyze the characteristics of the radiation noise in circumferential direction. The distance from the field point to the bearing axis was 210\u00a0mm. Assuming that the bearing operated steadily with the axial preload of 350\u00a0N, the directivity of sound field was explored at the rotational speed of 3000\u00a0r\u00a0min\u22121, 6000\u00a0r\u00a0min\u22121, 9000\u00a0r\u00a0min\u22121, 12,000\u00a0r\u00a0min\u22121, 15,000\u00a0r\u00a0min\u22121, and 18,000\u00a0r\u00a0min\u22121. The distribution of sound field in circumferential direction under different speeds is shown in Fig.\u00a010. As shown in Fig.\u00a010, the radiation noise increases with increasing the rotational speed and it has the maximum radiation noise at rated speed. In the circumferential direction with the same distance from field point to bearing axis, there is an uneven distribution of the sound field. The SPLs have a larger value at the upper left semicircle and at the right lower semicircle. A great impact noise caused by the ceramic ball and cage is in an angle range of 0\u201360\u00b0 (upper left semicircle), which can be called impact load zone" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002914_978-3-030-60330-4-Figure2.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002914_978-3-030-60330-4-Figure2.2-1.png", "caption": "Fig. 2.2 Kinematic scheme", "texts": [ " But if the rigidities of the diagonal springs in pairs are different, then there is a lattice with a chiral microstructure. Dynamical properties of such lattices were discussed, particularly, in Refs. [21, 22]. Eachparticle has three degrees of freedom,when itmoves in its plane: the displacement of the mass center of the particle with the number N = N(i, j) po oc m x i y (translational degrees of freedom ui, j and wi, j ) and the rotation with respect to the mass center (the rotational degree of freedom \u03d5i, j (Fig. 2.2). The kinetic energy of the particle N(i, j) equals Ti, j = M 2 ( u\u03072i, j + w\u03072 i, j ) + J 2 \u03d5\u03072 i, j . (2.1) Here, J = Md2/8 is the moment of inertia of the particle about the axis passing through its mass center. The upper dot denotes derivatives with respect to time. The displacements of the granules are supposed to be small in comparison with the sizes of the elementary cell of the lattice. The energy of each particle provided by deviation of the particle from the equilibrium state is determined by the strain energy of the springs connecting this particle with the six nearest neighbors in the lattice" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001165_s1068371214060042-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001165_s1068371214060042-Figure2-1.png", "caption": "Fig. 2. Slip power control scheme of a wound rotor induction generator.", "texts": [ " Moreover, in the proposed control technique, only a fixed reactive power (capac itive) compensation is done at the rotor side. There fore, no separate control of reactive power is needed and it improves the efficiency of the controller too. Moreover, with this control approach the induction generator operates at its optimum condition. There fore, in addition to higher efficiency and higher power factor, the line current decreases and more power is fed to grid. 2. CONVENTIONAL CONTROL SCHEME In a conventional slip power control scheme (Fig. 2), the speed of the wound rotor induction motor is controlled by varying the external resistance. When, this method is extended for the wound rotor induction generators, the characteristics of the machine change as shown in Fig. 3, which suits a wind turbine based power generation [14]. Figure 3 indicates the different operating points of WRIG under different wind speed. However, in this type of control, there are substantial losses in external resistors at high speed which reduces the overall efficiency of the system [16]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000486_s11771-019-4180-x-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000486_s11771-019-4180-x-Figure4-1.png", "caption": "Figure 4 Applied linear tip relief of \u0394=1.0: (a) Sketch; (b) Extruded area", "texts": [ " Step 4: Calculate angle c (\u00b0), a sin h c A r . Step 5: Find angle a (\u00b0), a=180\u00b0\u2212i. Step 6: Find angle b (\u00b0), b=180\u00b0\u2212a\u2212c. Step 7: Calculate rT2 (mm), rT2= 2 2 a a a a2 cos ,r L r L b rp 15 s, vUGV1 = (0.6,\u22121.4) m/s. The speed of the other UGVs are similar to the corresponding UGVs in Scenario 2. (a) Path planned by CS. (b) Path planned by BS.", "texts": [ " The presented results were obtained with the following parameters: NG = 3, NT = 3, PA initial = (\u221220,\u221230) m, hinitial = 0 rad, PG 1 (0) = (\u221235, 30) m, PG 2 (0) = (10, 30) m, PG 3 (0) = (0, 40) m, PT 1 = (30, 0) m, PT 2 = (\u221210,\u221230) m, PT 3 = (0, 0) m, vA = 10 m/s, rT i,com = 2.5 m (\u2200i \u2208 {1, 2, 3, . . . ,NT}), rG i,com = 2.5 m (\u2200i \u2208 {1, 2, 3, . . . ,NG}), Nloop = 2. In Scenario 1 (see Fig. 4), all UGVs are fixed. In Scenario 2 (see Fig. 5), all UGVs move at different speed vectors. In Scenario 3 (see Fig. 6), UGV1 has a change of its speed during 5 s< t < 15 s while the motion parameters of the other UGVs are same with those of the corresponding UGVs in Scenario 2. From the analysis of these solutions in the three scenarios, it can be seen that the paths generated by CS and BS can make the messen- ger UAV visit/revisit the communication neighbourhood of the moving UGVs periodically. Table 1 shows the tour lengths and the average computing time for visiting each task point or each UGV used by the above two methods in the three representative scenarios" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003689_iemdc47953.2021.9449496-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003689_iemdc47953.2021.9449496-Figure1-1.png", "caption": "Fig. 1. Diagram of different types of eccentricity fault. Os: stator center. Or: rotor center. Ow: center of rotation. (a) Healthy motor. (b) Static eccentricity. (c) Dynamic eccentricity. (d) Mixed eccentricity.", "texts": [ "ndex Terms\u2014Induction motor, fault detection, static eccentricity, motor current signature analysis I. Introduction Eccentricity is a type of motor fault caused by nonuniform air gap between the stator bore and the rotor. There are three types of motor eccentricity fault: the static eccentricity (SE), the dynamic eccentricity (DE), and the mixed eccentricity (ME). Figure 1 shows a diagram for the three types of eccentricity. Here, the point Ow is the center of rotation, Os is the center of the stator bore, and Or is the center of rotor. When the three points coincide, the motor is healthy with no eccentricity fault, as shown in Fig. 1a. In the case of SE, the points Or and Ow coincide, but are having an offset from the center of the stator bore The work was done when L. Zhou was with Mitsubishi Electric Research Laboratories. Corresponding Author: Bingnan Wang (bwang@merl.com) Os, as shown in Fig. 1b. When under DE fault, the rotor\u2019s center of rotation Ow is aligned with the stator center Os, but the rotor center Or is orbiting around the point Ow, as shown in Fig. 1c. A mixture of both static and dynamic eccentricity is called ME, where the points Or, Os, and Ow are not aligned with one another, as shown in Fig. 1d. Typically the SE fault of motors is created during the manufacturing process due to the ovality of the stator bore and the misalignment of bearings. The detection of SE fault of machines at an early stage is essential, as it can evolve into severe ME over the motor\u2019s operation due to the unbalanced magnetic pull, and lead to the breakdown of the machine. Throughout the years, a number of methods have been proposed for the SE fault detection [2], [5]\u2013[8]. The motor current signature analysis (MCSA) is one of the most widely used method due to its advantages of low-cost, reliability and simplicity, and the fact that no additional sensor is required to attach to the motor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000549_0731684419888588-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000549_0731684419888588-Figure6-1.png", "caption": "Figure 6. Axial compression failure mechanics model of 2D CTCS. (a) Elastic buckling, (b) plastic collapse, and (c) cell wall deformation.", "texts": [ " In addition, it is proportional to the thickness ratio and inversely proportional to cell angle and cell length. At the same time, the cell wall of the 2D cell in every plane has symmetric bending deformation. Thus, the four vertices of the 3D cell have clockwise displacement, making the cell have twist deformation, which is defined as twist angle per strain. While twist angle per strain is related to the cell angle, it is not affected by the thickness ratio and cell length. When the strain is large, the CTCS may failed. Figure 6 shows the failure analytical model of 2D CTCS under axial compression. The failure of NPR cell with elastic\u2013plastic material is mainly caused by elastic buckling and plastic collapse.33 For metal materials, the elastic buckling condition can be satisfied only when the relative density is very small, that is, the value of thickness ratio is less than 10 3.34 But the cell thickness ratio of CTCS ranges from 0.01 to 0.1, so the plastic collapse is the only failure mode of the 3D CTCS. In Figure 6(b), when the cell walls of CTCS tend to bend, the bending moment, created by force that is perpendicular to cell wall, may reach the fully plastic moment, giving two plastic hinges of each cell wall. This gives strain\u2013stress curve a plateau at the limited stress. FEA model of the 3D cellular structure under in-plane compression By building an FEA model, the in-plane compressive analysis of the 3D CTCS was carried out, which is shown in Figure 7(a). In the model, there are six cells in each direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001519_s12046-015-0427-x-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001519_s12046-015-0427-x-Figure1-1.png", "caption": "Figure 1. Geometry of angular (two point) contact.", "texts": [ " The key principle of this method is the replacement of the rolling elements under compression by nonlinear traction springs or nonlinear connector elements (in ABAQUS Standard) between the centers of curvature (Houpert 2001). Consequently, rolling elements under compression are modeled by nonlinear connector element working only in tension. This is physically and geometrically justified, since under load, the raceways shift toward each other and thus simultaneously altering the distance between their centers of curvature and compressing the rolling elements figure 1. By using this approach, it is possible to determine not only the equivalent contact load, but also the variation in the contact angle from the displacement of the curvature centers. A rolling element in a sector is defined by two nodes representing the centers of curvature of the raceways. Each node is linked to the corresponding opposite node by nonlinear traction springs or connector element (figure 2). The two zones of contact between the rolling element and raceway in a sector are modeled by rigid shells and coupled by rigid beam elements to the corresponding centers of curvature at two nodes materializing the contact ellipse, as shown in figure 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001083_1.4027926-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001083_1.4027926-Figure1-1.png", "caption": "Fig. 1 Compressor rotor experimental rig", "texts": [ " The ball bearing experiments are carried out, the vibration acceleration signals of the bearing house and casing are measured simultaneously, and the characteristics of the vibration signals are compared. (3) The fault features are extracted by means of the frequency spectrum analysis and the wavelet envelope analysis, and the sensitivities of the fault diagnoses based on the casing test signals and the bearing house signals are compared. 2.1 Compressor Rotor Experimental Rig. The aero-engine compressor rotor experimental rig was developed by the Design Institute for Aviation Products of Harbin Dong-an Engine Ltd. As shown in Fig. 1, in this experimental rig, the compressor rotor is driven by an electromotor and rotates in a prescribed rotating range. The motor, gear box, torque\u2013speed transducer, and compressor are connected in turn. The torque is transmitted by the connecting shaft and diaphragm coupling. After increasing the speed by using the gear box, driven by the motor, the compressor rotates in a counterclockwise direction as seen from the back. The gears and bearings in the gear box are lubricated by spattering oil, and are sealed by a cup packer" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000379_j.aca.2019.08.020-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000379_j.aca.2019.08.020-Figure1-1.png", "caption": "Fig. 1. Photograph of the magnetic actuation setup. The magnets are placed under the superhydrophobic coated glass substrate and are mounted on top of the computercontrolled XY stage. The superhydrophobic surface can be seen resting on a 3D-printed holder.", "texts": [ " After baking for 2 h at 70 C, the devices were cut on the borders, and inlets and outlets were punched for fluidic access. Then, the PDMS layer was bonded to a glass slide using a plasma cleaner (Harrick Plasma Inc., USA) to complete the microfluidic device (Fig. S1). Magnetic actuation was carried out using an XY Gantry linear translation stage (H2W Technologies). The coated glass slide was placed on a static 3D-printed frame with an array of permanent magnets directly mounted to the XY stage under the glass slide. A photograph of the magnetic actuation setup is shown in Fig. 1. The use of permanent magnets simplifies the operation of the system with reduced costs. For reproducible spectroscopic measurements it is necessary to pin droplets on specific spots on the SH surface. We formed SETs on the SH surface to achieve pinning of droplets. In the microfluidic application described here a hydrophilic substrate (e.g. glass) is covered with a superhydrophobic coating (Ultra Ever Dry\u2122). Hydrophilic regions are produced by scratching and removing the coating revealing the glass beneath" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002242_j.ijsolstr.2020.01.021-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002242_j.ijsolstr.2020.01.021-Figure11-1.png", "caption": "Fig. 11. The obtained final configurations corresponding to (a) 8 active members; (b) 16 active members.", "texts": [ " It can be seen that the kinematic paths driven y the 8 and 16 selected active members almost coincide in the rst 60 steps, indicating that the motion of the tensegrity tower t this stage is mainly driven by the same 8 active members. On he subsequent paths, the distance from x j to x m is reduced when ore active members are adopted. Obviously, it is mainly due to he contribution of the other 8 active members. It can be deter- ined that \u03c1 = 22.1% and 17.2% for the final configurations coresponding to the 8 and 16 selected active members, respectively, hich are shown in Fig. 11 . For the kinematic path driven by the 8 selected active mem- ers, the matrix F T U m in Eq. (15) is established for each configration obtained, and its minimum eigenvalue \u03bbf min is calculated. t can be found in Fig. 12 that the eigenvalues \u03bbf min corresponding o all the configurations are greater than zero, indicating that the a a o n s m s \u03bb fi \u03bb t s i c c a a d x m 7 7 p o F T g b b a t N m p t f i m g i c fi T 7 t c s 1 < b t fi m a ll obtained configurations are illustrated in Fig. 13 . Since they are ll greater than zero, the tower cannot maintain structural stability n the path due to the lack of prestress and kinematic indetermiacy, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000153_eeccis.2018.8692952-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000153_eeccis.2018.8692952-Figure3-1.png", "caption": "Fig. 3. Mesh topology", "texts": [], "surrounding_texts": [ "A processor type microcontroller which has been integrated with the Arduino IDE. ESP8266 is a low cost development board that consolidates GPIOs, 12C, UART, ADC, PWM and WiFi for rapid prototyping [6]. NodeMCU ESP8266 has a memory capacity of 4MB. Considering the implementation of consensus algorithms on hardware requires a large memory allocation, then the memory capacity of the NodeMCU ESP8266 is sufficient for the implementation of the minimum consensus." ] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure10.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure10.4-1.png", "caption": "Figure 10.4 Simplified diagram of table hydraulics", "texts": [], "surrounding_texts": [ "The column, guided on a dovetail slide, carries the wheelhead at its top end and contains the motor and belt drive to the wheel spindle. The column and wheelhead are raised and lowered through a screw and nut from a handwheel on the front of the machine. A telescopic guard is fitted to prevent grinding dust coming between the slide surfaces." ] }, { "image_filename": "designv11_22_0001270_imece2014-36664-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001270_imece2014-36664-Figure1-1.png", "caption": "Figure 1. A CAD Model used in EBAM experiments for temperature measurements", "texts": [ " The Arcam S12 EBAM machine at NASA\u2019s Marshall Space Flight Center was utilized to fabricate the test parts with Ti-6Al-4V powder material. The parts were fabricated each with a different speed function index: 20, 36, 50, and 65. Each part had a 60 mm by 5.25 mm cross-section, length by width (in X and Y) and was about 25 mm tall. The parts were spaced 1 mm apart from each other in order to fit all of them into the field of view of the camera. A CAD model of the parts in their arrangement for fabrication is shown in Figure 1. The EBAM machine uses a parameter setting called the speed function (SF) index to control the translational speed of the electron beam during the fabrication of the part. For each speed function value, though, the actual translational speed as well as the beam current will change along the build height. The build theme for the fabrication of Ti-6Al-4V parts with a layer thickness of 70 \u03bcm specifies the default speed function index value as 36. The information of the speed function algorithm can be found in Mahale [20] who noted that the beam speed dropped significantly during the fabrication of the first several millimeters of the part then gradually stabilized for the rest of the build" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001948_med.2016.7536007-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001948_med.2016.7536007-Figure1-1.png", "caption": "Fig. 1. Device used in the experiments for throwing balls.", "texts": [ " Although the dataset is artificially generated in the experiments it was used to predict the real trajectories. \u2022 A set of real trajectories acquired by the stereo camera system. The predictor tests were performed using leaveone-out method: each trajectory from the database was predicted using other examples. Structured prediction was implemented as a C++ program. It is executed purely on the CPU without any use of the GPU parallelization. A numerically controlled linear throwing device was used in the experiments (figure 1). It allows throwing balls with specified nominal velocity. Real launching velocity deviates from the nominal values (this is one of the mean reasons, why the trajectories differ from each other). The thrown balls intermits the light barrier, which is mounted in front of the throwing device. This event triggers the trajectory observation system, which consists of two cameras integrated into the stereo pair. The stereo system provides the measurement of object coordinates in space with a frequency of 100 frames per second" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001436_1350650113519611-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001436_1350650113519611-Figure5-1.png", "caption": "Figure 5. Experimental apparatus for measuring the bump foil structural stiffness.", "texts": [ " The air pressure obtained from the Reynolds equation was applied on the shell elements as distributed load. The bump foil was replaced by several spring elements and one end of spring elements was fixed on the base plate in all directions. In this paper, the spring coefficient of spring elements was experimentally determined by using the bump foil structural stiffness presented by Lee et al.13 Measurement of the bump foil structural stiffness The bump foil structural stiffness is needed for calculating the foil deformation, and was experimentally determined in this paper. Figure 5 shows the experimental apparatus for measuring the bump foil structural stiffness. The shaft was supported by two aerostatic journal bearings to move the shaft without friction in the vertical direction. Four bump foils were fixed at the position of the bearing base plate by spot welding as shown in Figure 1, and set at the upper end of the shaft. The static load was applied on four bump foils by an air cylinder installed at the lower end of the shaft. The deformation of the bump foils was measured by using two capacitive displacement sensors located at the upper end of the shaft" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002353_acs.analchem.9b05406-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002353_acs.analchem.9b05406-Figure1-1.png", "caption": "Figure 1. Schematic overviews of used setups: (A) EC-MS, (B) ECCE-MS, and (C) EC-CE-UV/vis. (a) Computer, (b) potentiostat, (c) time-of-flight mass spectrometer, (d) fused silica capillary, (e) flow cell with integrated SPCE, (f) syringe pump, (g) XZ-table with SPCE and buffer reservoirs, (h) HV source, (i) autosampler with SPCE and buffer reservoir, (j) UV detector.", "texts": [ "1% formic acid, Gradient: 0 min 100% A/0% B, 4 min 100% A/0% B, 8 min 2% A/98% B, 9 min 2% A/98% B, 9.1 min 100% A/ 0% B, 10.1 min 100% A/0% B; flow rate 0.4 mL min\u22121; injected volume: 2.5 \u03bcL. Ion source: AJS ESI; dry gas temperature 300 \u00b0C; dry gas flow 8 L min\u22121; nebulizer gas pressure 2.8 bar; sheath gas temperature 300 \u00b0C; sheath gas flow 10 L min\u22121. Collision induced dissociation of target species at 10 and 20 eV. Instrumental Setups and Parameters. A schematic overview of the instrumental setups is depicted in Figure 1. For voltammetric studies, 50 \u03bcL of 1 mM thymine solutions were applied onto screen-printed electrodes and potential scans were carried out at a scan rate of 50 mV s\u22121. As illustrated in Figure 1A, EC-MS measurements were carried out by coupling an electrochemical flow cell with integrated SPCE (DRPFLWCL, DropSens) to MS via a fused silica capillary (50 \u03bcm \u00d7 21 cm, Polymicro Technologies, AZ, USA), which was installed in the cell using a modified PEEK fitting and sleeve as described previously.18 The solutions were transported to the detector by a microinjection syringe pump (World Precision Instruments, Sarasota, FL, USA) equipped with a 1 mL glass syringe (Hamilton Company, Reno, NV, USA) at a flow rate of 16 \u03bcL min\u22121, corresponding to a fast transfer time of 1\u22122 s between generation at the electrode and detection in MS (experimentally determined by using ferrocene methanol as model system at an oxidation potential of 0.5 V). Due to this fast detection of products, the real-time response of the product composition while scanning a potential ramp from 0 to 2 V at 10 mV s\u22121 could be recorded. Solutions of 1 mM and https://dx.doi.org/10.1021/acs.analchem.9b05406 Anal. Chem. XXXX, XXX, XXX\u2212XXX B 0.1 mM thymine in 50 mM NH4OAc (pH 7.0) or 50 mM NH4HCO3 (pH 8.0) were used. EC-CE-MS (Figure 1B) and EC-CE-UV/vis measurements (Figure 1C) were carried out using laboratory-constructed CE systems, as described elsewhere.21,32 Samples were hydrodynamically injected into the separation capillary by placing the injection end of the capillary directly onto the working electrode surface during oxidation. The capillary tip was polished to an angle of 15\u00b0 at the injection end with a laboratory-constructed polishing machine using lapping foils with a grain size of 30 and 9 \u03bcm to ensure a reproducible injection position. For EC-CE-MS measurements, an oxidation potential of 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002563_j.robot.2020.103599-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002563_j.robot.2020.103599-Figure6-1.png", "caption": "Fig. 6. Diagram of static torque illustration of the bending module.", "texts": [ " (12)\u2013(14), \u2202V1/\u2206l and \u2202V2/\u2206l can be calculated as \u2202P1 \u2202\u2206l = \u2202P1 \u2202V1 \u2202V1 \u2202\u2206l = \u2212\u03b3 P1 l + \u2206l , (15) \u2202P2 \u2202\u2206l = \u2202P2 \u2202V2 \u2202V2 \u2202\u2206l = \u03b3 P2 l \u2212 \u2206l , (16) where \u03b3 = 1.4 for air. The stiffness is another important factor influencing the module\u2019s behavior and is defined as the ratio of the interaction force to the extension, which can be obtained by combining Eq. (11)\u2013(16) as K = \u2212 \u2202Fout \u2202\u2206l = [ Ar l + \u2206l \u2212 \u00b5\u2032c1(\u03b3 \u2212 1) ] P1 + [ Ar l \u2212 \u2206l \u2212 \u00b5\u2032c1(\u03b3 + 1) ] P2 + 4\u00b5tl4c3 c21 (l + \u2206l)5 (17) When pressure P1 and P2 are applied to Chamber 1 and Chamber 2 in the bending module with P1 > P2, a bending angle \u03b8 will be generated as shown in Fig. 6. The axial stretch ratio of Chamber 1 and Chamber 2 under this bending configuration can be written as \u03bb1,1 = h\u03b8 + l l , (18) \u03bb2,1 = 1, (19) where h is the distance between the centroid of the chamber cross-section and the middle plane of the bending module. The torque equilibrium can be formulated as \u03c4out = Ah(P1 \u2212 P2) \u2212 \u03c4f \u2212 \u03c4e, (20) where \u03c4out is the output torque, \u03c4f is the torque induced by friction, and \u03c4e is the torque generated by chamber stretch. The frictional torque \u03c4f can be calculated as \u03c4f = \u00b5\u2032h [ P1c1(l + h\u03b8 ) + P2c1(l \u2212 2h tan \u03b8 2 ) ] " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003296_j.measurement.2021.109056-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003296_j.measurement.2021.109056-Figure1-1.png", "caption": "Fig. 1. Sketch of a spiral centrifuge adapted from [11]. 1. Gear box reducer; 2. Ball bearing; 3. Ball bearing; 4. Solid outlet; 5. Outer drum rotor; 6. Inner spiral rotor; 7. Liquid outlet; 8. Fixing bolt; 9. Ball bearing; 10. Ball bearing; 11. Bearing housing; 12. Extension part of inner rotor; 13. Right end of drum.", "texts": [], "surrounding_texts": [ "Figs. 1 and 2 show schematic illustrations of a centrifugal separator system and a dual-rotor aeronautical turbine, respectively. For the centrifugal separator, the outer cylinder rotates at high speed to generate the centrifugal effect. Internally, a spiral rotates at a slightly greater rotation, pushing the solid parts towards the outlet. The speeds used for these rotors depend on the type of solid\u2013liquid mixture being filtered, but the difference in speed of rotation between these two rotors is often very small. For the dual-rotor aeronautical turbine system, the internal rotor is considered a low pressure rotor and the external rotor is considered a high pressure rotor. In some specific situations, the difference in speed of rotation between these two rotors can also be very small. Each of the rotors shown in Figs. 1 and 2 presents an independent unbalance and, even if they are within acceptable limits after manufacturing, when they work together, the sum of the existing unbalances can generate above ideal global levels of vibration, right after assembly or after a period of use due to wear. In experimental terms, tachometer and accelerometer sensors are used to capture the vibration signal at points of the mechanical system of greatest interest, usually on the bearing support structures. Subsequently, these acquired signals are traced with appropriate computational tools in order to obtain useful information for the design, maintenance or improvement of these systems. In the existence of beats frequency, the amplitudes of vibration add up and subtract each one depending on the moment of each force vector. When using a conventional vibration meter, the overall vibration value is indicated at each moment and, thus, the values fluctuate without a representative average that can be used for analysis. Therefore, in many practical situations it is not possible to perform a satisfactory vibration analysis and dynamic balancing using a conventional approach with conventional hardware and software." ] }, { "image_filename": "designv11_22_0000549_0731684419888588-Figure18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000549_0731684419888588-Figure18-1.png", "caption": "Figure 18. Three-dimensional chiral negative Poisson\u2019s ratio structure.21,22", "texts": [ " It can be seen from the table that the error of the effective Young\u2019s modulus of the numerical calculation and the theoretical analysis is relatively small, indicating that the finite element model is feasible. However, the results of the relative twist angle calculations between the numerical calculation and the theoretical analysis are quite different, since the theoretical analysis does not consider geometric nonlinearity and treats the ring as a rigid structure. However, the finite element model considers geometric nonlinearity and treats the ring as a structure with normal material. So there are big errors between them. Figure 18 shows the three different 3D chiral NPR structures. Table 7 shows the values of the elastic modulus, Poisson\u2019s ratio, and relative twist angle of the three 3D chiral structures. The structure of Figure 18 (a) proposed by Frenzel et al.31 and the structure of Figure 18(b) proposed by Fu et al.19 are isotropic on a macroscopic scale, which have the same elastic properties and NPR in the three directions in space. The 3D structure of Figure 18(c) proposed by Fu et al.21 has an anisotropy on a macroscopic scale, and the elastic modulus in the y direction is much larger than the elastic modulus in the x direction. In addition, although the Poisson\u2019s ratios in all directions are negative, the Poisson\u2019s ratio in the z direction is the smallest. Among these three structures, only the structure of Figure 18(a) has both NPR and compression torsion characteristic, and the elastic modulus is large, which opens up a new direction for the application field of multicellular structures. An in-plane compressive experiment was carried out to prove above analysis with a 2 2 2 CTCS sample. The sample is produced by 3D printing, and the material is chosen as Acrylonitrile Butadiene Styrene (ABS): elastic modulus is 2.2Gpa, and Poisson\u2019s ratio is 0.4. Due to the 3D printing method, the mechanical properties of the printed structure will be anisotropic" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003261_s10514-020-09959-0-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003261_s10514-020-09959-0-Figure2-1.png", "caption": "Fig. 2 Schematic of a skid-steered wheeled robot. The center of mass of the robot is assumed to be close to G \u2032, the black dot at coordinates (\u2212\u03b4x,\u2212\u03b4y) with respect to the geometric center of the robot", "texts": [ " The approach in this paper is based on deriving the rotational velocity of a skid-steered robot, assuming it is rotating at a point close to its center of mass. It is also assumed that the surface is firm with isotropic friction, and the platform is operating at relatively low speeds with no accelerations. Later it is experimentally verified that the derived expression is valid for an arbitrary instantaneous center of rotation. In this derivation it is assumed that the distance between the wheels of one side is a and the distance between both sides, i.e. the base width, is b (Fig. 2). The rotational speeds of the left and right side wheels are denoted by \u03a9L and \u03a9R . The corresponding translational speeds of a point on the surface of the wheels with respect to the axis of the wheel are represented by uL and uR . These speeds can be calculated using uL = r\u03a9L and uR = r\u03a9R , where r is radius of the wheels. The rotational speeds \u03a9L and \u03a9R are directly measured by the encoders. vxy is used to denote the velocity vector that is attached to the axis of wheel xy and is parallel to the plane of the wheel and the ground" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001300_1.4906967-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001300_1.4906967-Figure2-1.png", "caption": "FIG. 2. Illustrations of the flexible Ge p-i-n diode structure on a plastic substrate and the mechanical bending conditions applied on the diode.", "texts": [ " Patterning on the Ge layer is also different from that on the Si layer, since the Ge layer shows very slight color contrast between the doped and un-doped regions, thus align marks are necessary for every patterning process. After the undercut of the buried oxide layer in hydrofluoric acid (HF), the Ge nanomembrane are \u201cflip\u201d transferred to the plastic substrate. Finally, the metal interconnections and electrodes (stack of 40 nm Cr and 500 nm Au) are formed. Figure 1 shows the optical and microscopic images of the transferred Ge nanomembrane (light green color) and the finished Ge diodes on a plastic substrate. The device structure schematic is shown in Fig. 2. The Ge diodes have lateral p-intrinsic-n (p-i-n) structures, with two identical paralleled p-i-n channels. Ground-signal-ground (GSG) configurations are employed for the diode input and output ports. The example Ge diode under test has a a)Author to whom correspondence should be addressed. Electronic mail: gqin@tju.edu.cn 0003-6951/2015/106(4)/043504/4/$30.00 VC 2015 AIP Publishing LLC106, 043504-1 This article is copyrighted as indicated in the article. Reuse of AIP content is subject to the terms at: http://scitation", "5 mm (for convex bending) and 85 mm, 110.5 mm (for concave bending). On the 175 lm thick plastic substrate, the corresponding strains applied on the Ge diode are calculated to be 0.11%, 0.23%, 0.31%, 0.57% (tensile strain for convex bending) and 0.11%, 0.08% (compressive strain for concave bending), respectively. Transversal bendings (bending direction is perpendicular to the p-i-n channel) and longitudinal bendings (bending direction is paralleled to the p-i-n channel) are applied to the diode, as illustrated in Fig. 2. Figures 3(a) and 3(b) show the dc performance of the flexible Ge diode at forward bias under the transversal and longitudinal bending conditions, respectively. As shown in Fig. 3(a), under the transversal bending conditions, the forward current shows a slight degradation from larger tensile strains to larger compressive strains (i.e., from smaller convex bending radius to smaller concave bending radius). On the contrary, under the longitudinal bending conditions, the forward current improves as the diode is bended with larger tensile strains to larger compressive strains" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000588_9783527822201.ch15-Figure15.8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000588_9783527822201.ch15-Figure15.8-1.png", "caption": "Figure 15.8 Overview of the bending parameters for method A: (a) side view with typical curvature achieved by a deposition of an equal number of layers (n) with different thickness (h1, h2); (b) top view of the plotting area depicting relevant parameters for the model of the deposition; and (c) typical resulting structure.", "texts": [ " In particular, the bending can be obtained by two methods: (A) Continuous deposition: A deposition of the same number of layers k with different thicknesses h1 and h2 (Figure 15.7b) (B) Reversing deposition: A deposition of different number of layers n and m with the same thickness h (Figure 15.7c). Continuous Deposition (Method A) With this deposition strategy, the direction of the rotation is never inverted and the bending is obtained by tuning the layer height along the deposition circumference. A curvature (Figure 15.8) can be obtained by setting properly the layer thickness h at a generic point P \ud835\udefc along the circumference (with \ud835\udefc angle of P with x-axis): P \ud835\udefc = ( d 2 cos \ud835\udefc, d 2 sin \ud835\udefc ) (15.9) Considering \ud835\udf03 the growing direction with respect to the x-axis, in that position the deposition should be minimum (height h1) and in the opposite position (\ud835\udf03+\u03c0) the deposition should be maximum (height h2). The height of the layer can be calculated in a generic point P \ud835\udefc as a function of the angle \ud835\udefc: h \ud835\udefc = h1 \u2212 h2 d ( d 2 cos(\ud835\udefc \u2212 \ud835\udf03) \u2212 d 2 ) + h1 (15" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001681_ccdc.2014.6852993-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001681_ccdc.2014.6852993-Figure4-1.png", "caption": "Figure 4: Multi-body mass-spring model", "texts": [ " However, these above methods can not be well used to simulate the motion of a catheter/guide wire with the complex 3D vasculature model in real-time or near real-time [7], [8]. In this paper, a new method which is suitable for the real-time catheter/guide wire simulation is presented. When a catheter/guide wire is pushed, the external forces are applied to the two nodes of the link. The forces applied on the link push the tip at the point tipx(0) and pull the body at the point bodyx(0) , as shown in Fig. 4. How to solve the detection of the collision between a catheter/guide wire and a 3D blood vessel model is a major problem of the virtual minimally invasive surgery simulator. When the trainees steer the virtual catheter/guide wire by pushing, pulling and rotating at the proximal end of the catheter/guide wire, the collision detection algorithm is performed to avoid the catheter/guide wire from going through the 3D vascular model. In this paper, a real time collision detection algorithm is presented, which combines the median point triangle algorithm and the continuous collision detection algorithm [5]", " Three internal forces are spring force (Fspring), shear force (Fshear) and bending force (Fbending). Shear force is based on the hybird mass-spring model presented in [3]. Fexternal is composed of Finit , Fadd and Fenviroment . Finit represents the initial external force. Fadd represents the external force derived from a trainee. Fenviroment represents the forces derived from environment. When a catheter/guide wire is pushed, the external force is not applied to the last particle of the body, but to the two nodes of the link at each simulation loop, as shown in Fig. 4. The external force for the nodes of the link is given as follow: \u2022 First, we initialize the add force with the value Fadd and the minimum variation of the tip position with the value position(top). \u2022 At each simulation loop, \u0394tipx(0) is computed and compared with the position(top) (\u0394tipx(0) is obtained by subtracting the position of topx(0) at the moment of updata(i) from the moment updata(i+1)). 2014 26th Chinese Control and Decision Conference (CCDC) 4595 The Fadd increases at each frame until the \u0394tipx(0) is bigger than the position(tip) (Fadd + = F)", "tiplink \u2016 ) (3) where kspring.tiplink represents spring coefficient. Lspring.tiplink represents the distance between a tip node and the linkline. lspring.tiplink represents the relax distance between the ith node and the linkline. Uspring.tiplink represents the vector obtained by subtracting the position of the ith node from the position of corresponding point in the linkline. Fspring.tiplink is a rigid spring force to constrain an angular displacement (\u03b8 ) between the tip and the link, as shown in Fig. 4. Fspring(i) = Fs.spring(i)+Fa.spring(i) (4) Fs.spring(i) = \u2212ks.spring(Lxi\u2212xi\u22121 \u2212 l) \u00b7 ( Uxi\u2212xi\u22121 \u2016Uxi\u2212xi\u22121 \u2016 ) (5) where ks.spring represents spring coefficient. Lxi\u2212xi\u22121 represents the distance between the position of xi and the position of xi\u22121. l represents the relax distance between the position of xi and the position of xi\u22121. Uxi\u2212xi\u22121 represents the vector obtained by subtracting the position of xi from the position of xi\u22121. The force of Fs.spring connects the neighbouring nodes with a rigid spring", " The function of cross is defined as the cross-product operation (the crossproduct is given by the vector obtained by subtracting the position of xi from origin and the vector obtained by subtracting the position of xi\u22121 from origin) . Ux represents the vector thar is parallel to the X axis [10]-[12]. The force of Fbending(i) represents a 3D force applied at the xi, as shown in Fig. 7. As mentioned above, in order to accelerate the collision response, we set the velocity of the tip nodes to be Vtop, as shown in Fig. 4. The Vtop is defined as: Vtop = kv.top(bodyx(0)\u2212 topx(0)) (10) where kv.top represents the proportionality coefficient. When the guide is pulled, the external force applied on the last body particle. The motion of the body particles are computed starting from the the last particle to the tip particles at each loop, as shown in Fig. 4. The external force is given by: Fexternal = Finit +Fenviroment (11) 4596 2014 26th Chinese Control and Decision Conference (CCDC) The internal forces applied on the tip are two spring forces (Fspring.tiplink, Fspring). Fspring(i) = Fs.spring(i)+Fa.spring(i) (12) The internal force applied on the body is spring force (Fspring). Fspring(i) = Fs.spring(i) (13) When a rotation of angle \u03d5 is applied to the last particle of the catheter/guide wire body, the link and the tip are made to move in a circular as a whole" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000871_12.2185196-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000871_12.2185196-Figure1-1.png", "caption": "Figure 1. Quadrotor flight: A. Hover, B. Vertical climb, C. Lateral flight, D. Yaw maneuver", "texts": [ " The four rotors are configured into two clockwise and counter clockwise rotating pairs. The counter rotation of each pair serves to cancel out the net reaction moment when all rotors spin at the same angular velocity. In some quadrotors, the arrangement is that of a cross (+), where the primary directions of lateral flight (forward/reverser, left/right) are parallel to the rotor arms. In most other quadrotors, the configuration is that of an X, where the rotor arms are offset 45\u00b0 from these principle directions. Figure 1 depicts some of the main maneuvers of an X configured quadrotor, and how they are achieved by varying rotor thrust. Proc. of SPIE Vol. 9468 94680R-2 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 07/20/2015 Terms of Use: http://spiedl.org/terms When all rotors spin at the same speed and generate enough thrust to overcome the weight of the airframe, a stationary hover is achieved (Fig. 1A). When thrust on all rotors are increased equally, the quadrotor climbs vertically (Fig. 1B). Lateral flight is achieved by increasing two rotors\u2019 speed relative to their opposites, generating a net moment on the airframe which pitches the craft. The change in pitch vectors some of thrust into the horizontal plane, pushing the quadrotor forward (Fig. 1C). Orientation control is achieved by changing a given rotor pair\u2019s speed with respect to the second, resulting in net moment and causing the quadrotor to yaw about its center of gravity (Fig. 1D). Two basic approaches can been taken to model the quadrotor [4]: one using the Newton-Euler equations of motion [5, 6, 7, 8, 9] and the other the Lagrange-Euler equation [10, 11, 12]. The referenced sources are only a few examples of models employing either method. Many quadrotor models have been developed in recent years, the majority take the Newton-Euler approach, likely on account of its more intuitive appeal. The following model is based on the Newton-Euler approach, and follows the basic equations developed in [13]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000486_s11771-019-4180-x-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000486_s11771-019-4180-x-Figure5-1.png", "caption": "Figure 5 Applied parabolic tip relief of \u0394=1.0: (a) Sketch; (b) Extruded area", "texts": [ " Using the established tip relief procedure, linear tip modification was applied on the meshed pinion and gear teeth involute edges. The same value was used for both normalised amount and length of modification and it varied from 0.25 to 1.0 with increment of 0.25. Figure 4 shows the application of linear tip relief for a normalised tip relief value of 1.0. J. Cent. South Univ. (2019) 26: 2368\u22122378 2373 Similarly, parabolic tip relief was also applied using the same normalised amount and length of modification values used for linear tip relief (see Figure 5). The maximum frictionless contact stress of the non-modified gear set was recorded on the gear tooth subsurface with a value of 1227.5 MPa. Figure 6 shows the contact stress distribution which is higher close to the contact area and decreases away from it. Similarly, the maximum bending stress was also recorded on the gear tooth root. The stress can be seen to decrease with distance away from the gear tooth fillet area as shown in Figure 7. The non-modified gear set frictionless contact stress of 1227" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000861_j.phpro.2015.06.044-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000861_j.phpro.2015.06.044-Figure9-1.png", "caption": "Fig. 9.Photoelastic model Fig. 10. Flexure model loading (a) Radial Loading; (b) Axial Loading", "texts": [ " The magnitude of the relative retardation is given by the stress-optic law [7]: (1) Where is the induced retardation, Co is the stress-optic coefficient, t is the specimen thickness, 1 and 2 are the first and second principal stresses, respectively. It may be written as, 1 2 f N t (2) Where 1\u2013 2 is difference of principal stresses, f Co), depends upon the model material and the wavelength of light employed, and N = ( ), is the relative retardation of rays forming the pattern. The term N is also known as the isochromatic fringe order. Fig. 9 shows the photoelastic model for spiral and linear flexure bearings. The solid circular disc was made from photoelastic material. With the help of diametric compression method the material fringe value was calculated. The material fringe value for the model is f = 14.2. Fig. 10 shows the flexure model loaded in the polariscope arrangement. Radial and axial loading was applied on the spiral and linear flexures. The combined radial and axial loading was also applied and the stress distribution within the model was studied and high stress location points were identified" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003440_j.mechmachtheory.2021.104297-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003440_j.mechmachtheory.2021.104297-Figure3-1.png", "caption": "Fig. 3. The Hunt-type singular configuration of the 3\u20133 GSP: limbs 1 and 2 lie in the plane Q 1 Q 3 Q 5 of the moving platform.", "texts": [ " , 6 (70) where r i is the vector of the limb M i Q i ( i = 1, 2, \u2026, 6); b i is the vector of the i th limb arm relative to point O p ; and l i is the length of the i th limb, r i \u2261 [ x i y i z i ] = r Qi \u2212 r Mi = [ x Qi \u2212 x Mi y Qi \u2212 y Mi z Qi \u2212 z Mi ] = [ r 2 cos \u03b2i \u2212 r 1 cos \u03b1i r 2 sin \u03b2i \u2212 r 1 sin \u03b1i z i ] (71) l i = \u2016 r i \u2016 = \u221a x 2 i + y 2 i + z 2 i (72) \u03b1i and \u03b2 i are the angles defining the point of the i th limb applications in the base and moving platform, respectively. Let us consider a set of the 3\u20133 GSP configurations, in which the moving platform moves along the Z-axis and is tilted around the Y-axis at an angle of \u03c8 = 30 \u00b0 when the moving platform centroid O p is located on the Z -axis ( Fig. 3 ), i.e., x p = y p = 0, and the angles of rotation of the platform around the X- and Z-axis are \u03b8 = \u03d5 = 0. During the movement along the Z-axis the platform passes through the Hunt-type singular configuration [35] , which takes place when two limbs and the moving platform lie in the same plane. In the case shown in Fig. 3 , limbs 1 and 2 lie in a plane Q 1 Q 3 Q 5 of the moving platform. In this configuration, a sum of the moments relative to the line Q 3 Q 5 of six reactions produced in the limbs equals zero, and there is no possibility to counterbalance an active torque around the axis Q 3 Q 5 . The mechanism has an additional degree of freedom: an infinitesimal rotation around the axis Q 3 Q 5 can be carried out without any resistance whatsoever. The same is true for a force applied at point Q 1 ( Q 2 ) acting along the normal at the moving platform plane. Since lines O p Q 1 and M 1 M 2 lie in the same plane, they are intersected at a point Q . One can obtain from triangles OO p Q and M 1 Q 1 Q z 1 = ( r 1 / 2 ) tan \u03c8 (73) l 1 = l 2 = \u221a ( r 1 2 cos \u03c8 \u2212 r 2 )2 + 3 r 2 1 / 4 (74) where z p = OO p ; l 1 = M 1 Q 1 and l 2 = M 2 Q 2 are the limb lengths ( Fig. 3 ). Two other singular configurations in the finite-dimensional area are as follows: z 2 = \u2212( r 2 / 2 ) sin \u03c8 and z 3 = ( r 2 \u2212 r 1 ) tan (\u03c8/ 2) (75) The determinant of the Jacobian matrix vs. the Z -position of the moving platform, with r 1 = 0.7, r 2 = 0.25 m, and \u03c8 = \u03c0 /6, is shown in Fig. 4 , where the three singular configuration z 1 , z 2 and z 3, Eqs. (73 - 75 ) are fixed. Two more singular configurations take place out of the workspace with z \u2192 \u00b1\u221e . In Fig. 5 , the determinant vs. translations of the moving platform along X - and Z -axis is shown as the surface 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002888_s12221-020-1016-0-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002888_s12221-020-1016-0-Figure6-1.png", "caption": "Figure 6. Effect of bulge on forming depth (a) an extra force in plates when there was a bulge and (b) no extra force in plates without bulge.", "texts": [ " Thickness Toolpath Step size (mm) Depth (mm) Bulge Cracking 1 0.9 Outside-in 0.5 5 2 0.9 Outside-in 0.5 10 \u221a \u221a 3 0.9 Inside-out 1.8 10 4 0.9 Inside-out Change 10 5 0.9 Inside-out Change 15 6 0.9 Inside-out Change 20 \u221a 7 1.5 Outside-in 0.5 5 8 1.5 Outside-in 0.5 10 \u221a \u221a 9 1.5 Inside-out Change 10 10 1.5 Inside-out Change 15 11 1.5 Inside-out Change 20 12 1.5 Inside-out Change 25 \u221a found that the forming depth of the composite with the inside-out toolpath was larger than that with the outside-in toolpath, as shown in Table 4. As shown in Figure 6(a), when there is bulge, the tool would bear an extra force generated by bulge. And thus the plate would bear higher force when the bulge was larger with the deeper processing depth. Thus, the thickness of plate would thin with the processing depth. That\u2019s to say, the plate would crack earlier and the tool would wear out more easily. On the contrary, the thickness was uniform without bulge, as shown in Figure 6(b). The experiments continued normally until the plate was broken. In the three-tier-plate clamping method, the forming depth of the upper iron plate had a positive impact on the forming depth of composite sheet. Thus, it was necessary to investigate the effect of toolpath on forming depth of the iron upper plate. Cosine law model was commonly used in the thickness prediction of metal sheet under forming [20]. In the three-tier-plate clamping method, the thickness of the upper iron plate accorded with the cosine law model" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000450_j.eml.2019.100554-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000450_j.eml.2019.100554-Figure1-1.png", "caption": "Fig. 1. (a-b) 3D freeze-printed structures (c) solid\u2013liquid interface during freeze-printing (d) angle dependent filament diameters. (e) Oxide encapsulating the liquid metal with expected stresses at the nozzle exit (f) free body diagram demonstrating various forces acting on the oxide skin.", "texts": [ " This oxide skin acts as an elastic solid shell encapsulating the liquid metal inside, aiding in shaping the liquid metal droplets into filaments [19]. Utilizing this phenomenon, numerous research groups have demonstrated various applications and techniques to additively manufacture liquid metallic structures, most notably with room temperature liquid alloys such as eutectic alloy of gallium and indium (EGaIn) [14,20]. Recently, our research group has demonstrated free form direct-ink extrusion of EGaIn filaments in complex non-planar 3D geometries (Fig. 1(a-b)). These structures were realized by utilizing the temporary structural stability of oxide skin (primarily consisting of gallium oxide, 3\u20135 nm in thickness), along with freezing [21,22] as shown in Fig. 1(c). During these experimental studies, we observed some interesting phenomena pertaining to liquid metal extrusion: At a constant printing pressure and printing speed, the filament diameters decreased with decreasing printing angle (as shown in Fig. 1(d)), indicating the variation of flow rate. Moreover, within the stable filamentary extrusion region, the filament diameters did not seem to be a function of printing speeds. A complete detailed experimental study involving these observations is presented elsewhere [21]. These observations indicate that extrusion of liquid metals is dominated by the interfacial mechanics of the oxide skin rather than the bulk properties of the liquid metal. Evidently, it is critical to understand the evolution of interfacial stresses on the oxide skin during direct-ink extrusion of liquid metals to obtain a holistic picture of the printing process", " In that, first a droplet of the metal at the nozzle exit is created using a high-pressure pulse. This droplet is brought to contact with the substrate surface and the nozzle is then translated in the printing direction. As the liquid metal extrudes out of the nozzle under the influence of an applied pressure, a self-passivating oxide skin instantaneously forms on the surface. This self-passivating skin helps maintain filamentary shape as the nozzle translates in arbitrary directions defined by angle \u03b8 shown in Fig. 1(e). As the extrusion proceeds, material flow is achieved through the continuous yielding of the oxide skin under the tensile stress it experiences between the nozzle and the substrate as the adhesion between the initially formed droplet and the substrate serves as a \u2018\u2018clamping\u2019\u2019 effect [20,21]. Based on these findings, the oxide skin is expected to be stressed in three directions as shown: An axial stress (\u03c3zz) will develop along the printing direction along which the nozzle translates. Simultaneously, nozzle walls induce a shear stress (\u03c4s) on the extruding material", " Finally, similar to thin walled cylinders, the liquid core applies an internal pressure on the oxide shell, resulting a circumferential stress (\u03c3\u03b8\u03b8 ) on the oxide skin. It should be noted that \u03c3zz, \u03c3\u03b8\u03b8 and \u03c4s represent surface stresses with the unit of N/m rather than bulk stresses in N/m [2], given that the oxide skin thickness is nanometric [28] , thus negligibly small as compared to the filament diameter. The equations governing these stresses can be derived using the free-body diagram of the printed filament given in Fig. 1(f). Here, Va and Ha are vertical and horizontal reaction forces which will be measured, Dfil is the printed filament diameter, \u03b8 is the printing angle (0 for horizontal and 90\u25e6for vertical printing), w is the printed filament weight, Hn and Rn are the horizontal and vertical forces carried by the oxide skin at the top cross-section, Pn is the pressure force due to the hydrostatic pressure, Pe, at the nozzle exit. A momentum balance considering the fact that there is no flow inside the filament and the printed filament is not accelerating in any way boils to problem down to a simple statics analysis, as a result of which the three stress components can be written as: \u03c3zz = PeDfil 4 + VaSin(\u03b8 ) \u03c0Dfil + HaCos(\u03b8 ) \u03c0Dfil + wSin(\u03b8 ) \u03c0Dfil (1) \u03c3\u03b8\u03b8 = PeDfil 2 for cylinder (2", " Considering the Newtonian flow [26] of the liquid metal inside the nozzles Pe can be calculated through the Bernoulli principle by Pe = Pd + \u03c1gz \u2212 Ploss (3) where Pd is the dispensing pressure, \u03c1 is density of the liquid metal, g is gravitational acceleration, z is the liquid height inside the syringe and Ploss is the viscous pressure loss inside the syringe. Here, Pd is prescribed by the user using the dispensing system, z is measured prior to the experiments and Ploss is calculated through the flow rate and the nozzle diameter as described in our earlier work [21]. As shown in Fig. 1(e-f), the extruding material experiences a shear stress given by, \u03c4s = Fsp \u03c0Dproj (4) here Fsp is the force along the shear plane and Dproj is the filament diameter projected on the shear plane. The derivations for Eqs. (1)\u2013(4) is provided in the supporting information (S.1). In order to study oxide skin stresses, we used EGaIn (Gallium (75%) and Indium (25%), m.p. 15 \u25e6C). The EGaIn was selected due to two main reasons: (1) Since EGaIn\u2019s melting point is below room temperature, no liquid\u2013solid phase change occurs during printing at room temperature" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002348_j.mechmachtheory.2020.103871-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002348_j.mechmachtheory.2020.103871-Figure15-1.png", "caption": "Fig. 15. Contact simulation of circumferential stress of flexspline.", "texts": [ " Besides the influence of the tooth, there are many factors that need to be considered, such as the size and assembling position of the WG, the coning of the originally cylindrical flexspline, the flexion of the flexspline, and the Poisson\u2019 ratio, etc. This section mainly discusses the influence of the flexspline\u2019s structure, so to minimize the effect of the WG, the width of it is set to 1 mm. And since the flange structure is fixed and not deformed during operation, the established finite element models of flexsplines do not include the flange structure and ignore all of the fillets and chamfers except the bottom fillet of the cups. Fig. 15 shows the nephogram of circumferential stress results obtained by the contact analysis of the flexsplines and the elliptic cam WGs which parameters are in Table 7 . In this simulation, symmetrical constraints are applied on the 1/2 model, and the WG is positioned at the middle section of TR to ensure the maximum radial displacement here achieves w 0 . The material is set with elastic modulus E = 196 GPa and Poisson\u2019s ratio \u03bd= 0.3. Comparing Fig. 14 a with Fig. 15 a, it can be seen that the maximum circumferential stress (226.2 MPa) of the cup-shaped flexspline (Example-1) occurs near the front section of TR on the major axis, which increases about 27.6% as much as the stress (177.3 MPa) at the same position of the simple TR. The minimum stress ( \u2212177.5 MPa) occurs near the front section of TR on the minor axis is 16.5% larger (in absolute values) than the stress ( \u2212152.4 MPa) at the same position of the simple TR. Comparing Fig. 14 b with Fig. 15 b, the maximum circumferential stress (120.6 MPa) of the flexspline (Example-2) occurs at the middle section of TR near the major axis, which increases about 35.5% as much as the stress (89.0 MPa) at the same position of the simple TR. The minimum stress ( \u221296.2 MPa) occurs at the minor axis, is 24.6% larger (in absolute values) than the stress at the same position of the simple TR. The Poisson effect, which makes flexspline changes from 2D stress state to 3D stress state, is one of the causes that the maximum stress of flexspline is much larger than that of TR" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002408_s11012-020-01162-w-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002408_s11012-020-01162-w-Figure10-1.png", "caption": "Fig. 10 Bond graph of the aerodynamic block including the aerodynamic forces and their transformations", "texts": [ " The modules of the modulated transformers are embedded from the kinematic relations (39) and (41). Figure 9 depicts the related bond graph to the torsional mode shapes of the wings, which is constructed according to relation (66). As can be observed in the figure, three torsional modes have been considered for the wing. The modulus of transformers, which are equal to the torsional mode shapes have been calculated from Eq. (65). T is obtained from Eq. (56), and Twing is the wing elastic torsion, which is equal to the summation of the torsional mode shapes of the wing. Figure 10 indicates the way of producing the aerodynamic lift and drag forces of the wing both sections, illustrating them in the coordinate planes, and generating the forces and moments transmitted from the wings to the body. Vp1 and Vp2 are the linear velocities of the wing first and second sections, Vx and Vz are the bird vertical and horizontal velocities, and TW is the wing torsional torque, which construct the inputs of the aerodynamic block. Finally, Fx, Fy and Fz forces, M u, M h and Mw moments, and the aerodynamic torsion distribution on the wing T constitute the outputs of the aerodynamic block" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003676_s00170-021-07413-8-Figure31-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003676_s00170-021-07413-8-Figure31-1.png", "caption": "Fig. 31 Wrinkling defect in the sheet as the equivalent of columns buckling", "texts": [ " According to the figure, the thickness of the sheet has a significant impact on reducing and even preventing the wrinkles in the wall of the cone. Wrinkling in the sidewall depends on the ratio of the sheet free area between die and punch and blank thickness. As a result, the formation of wrinkles in the wall, because of the large aspect ratio of wall area to thickness, is more efficient. So, by increasing the thickness, from 1 to 1.5, and then 2mm, this ratio is reduced to 2 and will result in a reduction of wrinkling. In fact, if a narrow band of the cone wall was assumed as an equivalent column (Fig. 31) that is pressurized from the two sides (like hoop compressive stress in wall of cone), by increasing the thickness, the possibility of the column buckling decreases (like wrinkling in the wall of cone). Besides, maximum plastic strain variation on the wall of the formed part can provide a good assessment of the severity of wrinkles. In this regard, as shown in Fig. 30, increasing the thickness of the sheet leads to a significant reduction in the maximum plastic strain difference in the wall of the piece" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002286_j.cagd.2020.101826-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002286_j.cagd.2020.101826-Figure4-1.png", "caption": "Fig. 4. Interpreting lines as parallel in (a) parallel and (b) perspective projection.", "texts": [ " 3, we have shifted the position of the viewpoint, from the accidental position O, and then displayed the resulting views. The human brain is much better at spotting gaps (or overlaps) than a small curvature of an edge, so in S the pseudo-connection is clearly exposed, whereas for S\u0302 the deception still works. Yet not explicitly mentioned in the book by Hoffman (2000), we could add another rule for generic views: (3) If two straight lines look parallel in 2D, then they are parallel in 3D. In parallel projection (i.e., with O at infinity), if two edges are parallel in 2D, then so are they interpreted in 3D (Fig. 4a). Perspective projection is trickier, as it does not preserve parallelism, so we interpret several edges as parallel when they meet at a common vanishing point V outside the drawing (Fig. 5b). Also by grouping lines converging to vanishing points (Company et al., 2014), automated reconstruction methods interpret sketches of polyhedra in perspective projection. Sugihara (2018) has resorted to the violation of this rule in his award-winning impossible motion objects. We will also use this technique later to materialize the Renault logo (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003734_aer.2021.53-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003734_aer.2021.53-Figure16-1.png", "caption": "Figure 16. Controller integration and flight tests of the Ultra Stick 25e RC aircraft.", "texts": [ " HIL simulations are performed with the same reference variables tested for the SIL simulations with the backstepping and PID gains unchanged. Figure 14 validates the real-time implementation; the tracking achieved with the microcontroller is accurate and virtually identical to that obtained in the SIL case. The commands for this simulation are represented in Fig. 15, showing an excellent matching with the commands from the SIL. The controller has been integrated onto an Ultra Stick 25e RC aircraft model (Fig. 16), and preliminary validation flight tests have been conducted(39). The upgrade of the controller is under consideration. The initial control strategy was deliberately kept simple to demonstrate the feasibility of the real-time implementation and perform flight tests. Adaptive backstepping and substitution of the PIDs with more advanced laws are the changes investigated. 7.0 CONCLUSIONS An autopilot configuration combining non-linear control with the traditional PID technique is presented. The backstepping controller is employed to stabilise fast inner-loop variables characterising the aircraft attitude and aerodynamic angles, while PID gains control more slowly changing navigation variables" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001798_tcst.2016.2558538-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001798_tcst.2016.2558538-Figure3-1.png", "caption": "Fig. 3. (a) Definition of LOS frame B and the coordinate system (r, \u03b8 ) where the pursuer P is pursuing the target T . (b) Decomposition of the acceleration into radial component \u03bc and normal component \u03c3 in the LOS frame.", "texts": [ " Section III introduces the basic specifications of the experimental testbed. Section IV derives the bioinspired pursuit law using Lyapunov-based control and compares it with existing pursuit laws from missile guidance. Section V presents the experimental results using the hovercraft testbed. Section VI summarizes this paper and ongoing and future work. Consider the following formulation of the pursuit problem as a planar system of two point particles with unit mass. Let T and P denote the target and the pursuer, respectively. Fig. 3 depicts the relevant reference frames and coordinates, which includes the inertial frame I (O, x\u0302, y\u0302, z\u0302), the LOS frame B (P, e\u0302r , e\u0302\u03b8 , z\u0302), the relative position vector r = rT/O \u2212 rP/O , the range r = \u2016r\u2016, and the LOS angle \u03b8 , where cos \u03b8 = e\u0302r \u00b7 x\u0302. The inertial kinematics of the two-particle system expressed as components in frame B are [18][Id2 dt2 r ] B = [ r\u0308 \u2212 r \u03b8\u03072 2r\u0307 \u03b8\u0307 + r \u03b8\u0308 ] B = [ \u03bcT \u2212 \u03bcP \u03c3T \u2212 \u03c3P ] B (1) where \u03bcT and \u03c3T (resp. \u03bcP and \u03c3P ) denote the radial and normal components of the acceleration of T (resp" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002704_lra.2020.3013860-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002704_lra.2020.3013860-Figure7-1.png", "caption": "Fig. 7. Experimental setup: Two 1-DoF rotational devices with a hard wall contact.", "texts": [ " The proposed Proxy-based controller uses a common Lyapunov function to prove that the main-proxy subsystem is asymptotically stable for all the three switching instances. The secondary controller, using a stable trajectory following controller, follows the delayed main reference, and therefore the main, proxy and secondary positions converge. The human operator, while interacting with the MD, injects energy into the main-proxy coupling. However, when the operator releases the MD, the proposed controller decreases the energy monotonously due to dissipation by the inherent physical damping of the MD and the adaptive damping element of the PC. Fig. 7 shows two 1-DoF rotational devices used as the teleoperation setup. Both have a sampling rate of 1 kHz. A P-P control architecture with coupling gain of km = 1.5 N/mm, bmc = 0.01 Authorized licensed use limited to: University of New South Wales. Downloaded on August 23,2020 at 12:45:19 UTC from IEEE Xplore. Restrictions apply. Ns/mm on the main side and, ks = 1.5 N/mm, bsc = 0.01 Ns/mm on the secondary side was implemented. The gains were tuned such that at no-delay the energy generated by discretization was dissipated and therefore the teleoperator was passive", " the PCs act with a variable damping on the velocity signals (admittance type PCs) that exit the TDPNs in direction to the coupling controllers.The major drawback of admittance type PCs is a position drift that appears after the integration of a varied velocity signal. Drift compensation methods cannot compensate the drift instantaneously but only when the energy in the system and thus the physical coupling situation allows. Especially in P-P architectures critical coupling configurations can appear due to position drift. In the teleoperation experiment (compare Fig. 7) at Trt = 30 ms round-trip delay presented in Fig. 12(a), the MD commands a motion of theSD. At t = 2.5 s, a position drift is visible since the secondary reference position (xmd) does not match with the MD position (xm) and the main reference position (xsd) does not match with the SD position (xs). Occasionally, the drift equals for both PCs such that the coupling controllers find a suitable configuration. In contrast, the experiment displayed in Fig. 12(b) (Trt = 100 ms), the drift in xsd is higher than the drift in xmd" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure9.37-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure9.37-1.png", "caption": "Figure 9.37 Thread indicator dial", "texts": [ " To revert to a normal operating condition, the connecting rod is unclamped and the clamp bracket removed and, since the complete attachment moves with the carriage, the cross-slide can then be used in the normal way. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 9 142 9 Turning The position of engagement of the split nut on the leadscrew for each cut is important in order that the tool will travel along the same path as the previous cut. To achieve this accuracy of engagement, a thread indicator dial is fitted at the end of the apron, Fig.\u00a09.37. The dial is mounted on a spindle at the opposite end of which is a gear in mesh with the leadscrew. These gears are interchangeable, are stored on the spindle, and are selected by referring to a chart on the unit. They are arranged to give a multiple of the pitch required, relative to the 6 mm pitch of the leadscrew. The thread now standardised in British industry is the ISO metric thread, ISO being the International Organization for Standardization. Terminology of this thread is shown in Appendix 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000126_aer.2018.149-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000126_aer.2018.149-Figure1-1.png", "caption": "Figure 1. (Colour online) Sectional drawing of Rig 250 Build 6.", "texts": [ " In doing so, the mistuning strength is taken into account, which enables more realistic computations of forced response magnifications caused by mistuning. Petrov and Ewins(5) used optimisation algorithms to find the worst forced response of bladed disks in terms of academic studies. Similar objectives have been addressed in Refs 6 and 7. However, the majority of publications were considering measured or preset mistuning patterns, yielding magnification factors from 1 to hardly greater than 2 as exemplarily reported in Ref. 8. In this paper, real blade mistuning of an HPC test blisk (Rotor 2 of Rig 250; Fig. 1) is experimentally determined in terms of blade-to-blade frequency variations in order to use this data for updating structural models as close to reality as possible. Since every blade of the R2 blisk is applied with s/g it is a main objective to quantify the contribution of the s/g instrumentation and in addition the effect of assembling the test compressor on frequency variations. Reduced order models based on the subset of nominal system modes (SNM)(9) are employed in which mistuning is quantified by stiffness variations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001362_s11581-015-1426-y-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001362_s11581-015-1426-y-Figure1-1.png", "caption": "Fig. 1 aUnit PAFC experimental set-up with electrolyte electrode assembly. A cross-sectional view of graphite grooved plates is also shown as an extension. b The aqueous phosphoric acid electrolysis cell (Colour figure online)", "texts": [ " In this paper, we have reported in situ, concentrationdependent proton conductivity of aqueous phosphoric acid by constructing a unit PAFC and an electrolysis cell. The variation of conductivity at different PAFC cell temperatures has been studied also. The experimental results obtained from different cells have been compared and analyzed by theoretical simulation and ab initio calculation using density functional theory of phosphoric acid which has not yet been reported. Proton conductivity measurement using unit PAFC The schematic representation of unit PAFC experiment is shown in Fig. 1a. The unit PAFC consists of anode, cathode and electrolyte. Two electrodes were composed of thin layer of platinum (20 % Pt/C) deposited on to carbon plate. A glass mat soaked in aqueous phosphoric acid was used as an electrolyte. This electrolyte electrode assembly was placed between two graphite grooved plates as seen in Fig. 1a. Pure H2 was passed through a humidifier to moisturize the gas and then connected to the anode through the graphite grooved plates. On the other hand, pure O2 was directly connected to the inlet of cathode through the graphite grooved plates. The outlets of these graphite grooved plates were connected to adsorbers to adsorb the gained moisture. The cross section of the graphite grooved plate for maximizing the gas flow is also shown as an extended figure in Fig. 1a. Two stainless steel plates were placed at two ends as shown in Fig. 1a and used as current collectors. A heating plate was inserted after insulating electrically on the lower current collector to maintain the cell temperature. The total arrangement was kept compact by two pusher plates. Throughout all the experiments, the inflows of H2 and O2 were maintained at 100 and 50 cc/min, respectively, as monitored by rotameter. The terminals of cathode and anode were connected to impedance analyzer (Metrohm Autolab, Model PGSTAT 302 N). Measurements were performed at cell temperature from 85 to 130 \u00b0C and at fixed humidifier temperature 70 \u00b0C with a fluctuation of \u00b10", " The initial concentration of aqueous phosphoric acid was 88 % as obtained from Merck, India. The concentration of phosphoric acid was lowered by adding appropriate amount of deionized water. Following the conductivitymeasurements, the acid concentration was determined using a pH meter to detect the first equivalent point. Proton conductivity measurement using phosphoric acid electrolysis cell To measure the proton conductivity directly, we have chosen an electrolysis cell made of glass with two holes on the upper side as seen in Fig. 1b. Sufficient amount of aqueous phosphoric acid was taken in the electrolysis cell so that the two Pt disc electrodes were dipped into the acid through the holes of the electrolysis cell and the Pt disc electrodes were connected to the two probes of the impedance analyzer using Pt wires. The set-up was placed on the heating plate as seen in Fig. 1b. The conductivity measurement was performed after 15 min of temperature stabilization. Geometry optimizations were performed using Jaguar 8.5 implemented in Schr\u00f6dinger material suite [18]. Full optimizations over all degrees of freedom were carried out by conjugate gradient method [19] without symmetry constraints initially invoking in Hartree-Fock theory with the 6-31G** split valence basis set [20]. The resulting structures were refined thereafter with density functional theory employing Becke\u2019s threeparameter hybrid functional (B3LYP) [21, 22] and the same basis set", " Water evaporation and condensation of phosphoric acid above 100 \u00b0C may be a problem in proton conductivity measurement. In our experiments, the phosphoric acid is used in its concentrated form (88 and 80 %). Thus, the use of concentrated acid minimizes the water vapour pressure, and hence, water management in the cell is not difficult to achieve [8, 9]. Additionally, if there is a chance for water Type of experiment R\u03a9 (\u03a9) (\u00b10.01) Rct (\u03a9) (\u00b11) Diffusion time (ms) Qg (Fs n\u22121)\u00d710\u22129 n evaporation, the inlets of the experimental set-up (see Fig. 1b) are made in such a fashion that the tendency gets reduced. We have also checked the concentration of H3PO4 after the proton conductivity measurement at temperature 130 \u00b0C, and we have found no significant variation in concentration. On the other hand, the proton conductivity measurement has performed readily as described in the BExperimental procedure^ section. Consequently, the linear response of the proton conductivity indicates that the protons follow the Grotthuss mechanism. The Grotthuss mechanism is a common parlance for describing anomalous transport processes in aqueous acidic and basic solutions [14]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001552_taes.2015.140340-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001552_taes.2015.140340-Figure1-1.png", "caption": "Fig. 1. Geometric view of Doppler tracking.", "texts": [ " Section III is devoted to observability analysis of a constant velocity motion from a single Doppler sensor. Section IV presents the novel hierarchical fusion dual-stage filter (HF-DSF) and peer-to-peer DSF (P2P-DSF) for DOT, whose performance is evaluated in Section V via simulation experiments. Finally, Section VI ends the paper with concluding remarks and perspectives for future work. A Doppler sensor measures the frequency shift fd of a received target echo. By the Doppler effect, such a shift is proportional to the radial speed \u03c1\u0307 (see Fig. 1) according to the approximate relationship fd = fT \u2212 fR \u223c= 2 \u03bb \u03c1\u0307 = 2 \u03bb \u03bd cos \u03b1 (1) where: fT and fR denote the transmitted and received signal frequencies, respectively; \u03bb = c/fT is the transmitted signal wavelength, c being the light speed; \u03c1\u0307 = \u03bd cos \u03b1, \u03bd being the target speed; and \u03b1 the angle between the target direction and the sensor-target line, referred to in the sequel as the relative heading. Let us consider a target, moving in the (\u03be, \u03b7) plane, with kinematic state xt = [\u03bet , \u03be\u0307t , \u03b7t , \u03b7\u0307t ] evolving in time according to a discrete-time motion model xt+1 = f(xt ) + wt ", " In order to decompose the kinematic state x with respect to its observability properties from the Doppler measurements yt , t = 0, 1, . . . , one can resort to geometrical considerations and define a suitable transformed set of coordinates. To this end, let \u03c1t = |pt | be the range; \u03bdt = |p\u0307t | be the speed; \u03b8t = atan2(\u03b7t , \u03bet ) be the azimuth; and \u03b1t = atan2(\u03b7\u0307t , \u03be\u0307t ) \u2212 atan2(\u03b7t , \u03bet ) be the difference of the heading angle minus the azimuth, i.e., the relative heading. The geometrical meaning of such transformed coordinates is clearly depicted in Fig. 1. Next, the aim is to reexpress the constant velocity motion model in the transformed state vector z = [\u03c1, \u03bd, \u03b1, \u03b8] instead of the original state vector x. With this respect, from (6) and the definitions of the transformed coordinates, the following result can be readily stated. PROPOSITION 1 Consider the transformed set of coordinates z = [\u03c1, \u03bd, \u03b1, \u03b8] , then the state equations in the transformed state variables turn out to be\u23a7\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23a9 \u03c1t+1 = [ \u03c12 t + T 2 s \u03bd2 t + 2Ts\u03c1t\u03bdt cos \u03b1t ] 1 2 (7) \u03bdt+1 = \u03bdt (8) \u03b1t+1 = \u03b1t \u2212 atan2 (Ts\u03bdt sin \u03b1t , \u03c1t + Ts\u03bdt cos \u03b1t ) (9) \u03b8t+1 = \u03b8t + atan2 (Ts\u03bdt sin \u03b1t , \u03c1t + Ts\u03bdt cos \u03b1t ) (10) and the Doppler observation equation takes the form yt = \u03bdt cos \u03b1t " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000197_00450618.2019.1609088-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000197_00450618.2019.1609088-Figure6-1.png", "caption": "Figure 6. MJM printer.26", "texts": [ "16,20,21 Electron beam melting (EBM) printers are similar except they use a high voltage electron laser beam to melt metal powders. To avoid oxidation of metal parts the object is printed with the print bed in a high vacuum chamber.23 MultiJet Modelling (MJM) printers use inkjet print heads to deposit layers of liquid photopolymers onto the print bed, much like a 2D printer. Each layer is rapidly cured with UV light.16,24,25 MJM printers have a distinct advantage because the inkjet print heads can print using multiple materials at the same time 24. Figure 6 illustrates the MJM process. Figure 5. Example of SLM.22 Figure 4. SLA process.18 Three-dimensional printable materials can be categorized as polymers (filaments, powders, resins), metals or bio-inks.4 The printable materials are machine specific depending on the technological process being used and the object being printed.27 3.1. Polymer filaments Polymers are the most common material used in consumer 3D printing. These polymers typically lack functional strength and load-bearing capacity 27,28 and generate lowperformance objects" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000297_physrevapplied.12.014025-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000297_physrevapplied.12.014025-Figure2-1.png", "caption": "FIG. 2. Planar V-shaped structures with 120\u25e6 central angle together with their principal rotation axes {e1, e2, e3}: (a) slim structure with rectangular cross section (height-to-width aspect ratio h:w = 1:3); (b) chubby structure (square cross section). The structures are magnetized in-plane and the red arrows stand for the magnetic dipolar moment m.", "texts": [ " Taking a scalar product on both sides of this equation with = \u03c9z\u0302, we readily obtain it in a compact covariant form as Uz \u03c9 = \u0302 \u00b7 Ch \u00b7 \u0302 , (13) where Ch is a dimensionless chirality matrix given by the symmetric part of (1/ )G \u00b7 F\u22121, with the characteristic length and \u0302 = /\u03c9 = z\u0302 the normalized (unit) angular velocity. It is most convenient to write the rhs of Eq. (13) in the body frame, whereas Ch is fixed and \u0302 is expressed via the Euler angles as in Eq. (9). Note that Ch (in contrast to G) is independent of the choice of coordinate origin. 014025-3 Under the rotation of the coordinate frame, it transforms as a (symmetric) pseudotensor. Applying Eq. (13) to the symmetric V-shaped object (see Fig. 2), for which Ch has a single pair of identical nonzero off-diagonal entries, is straightforward. For such structures, the easy rotation axis e3 (corresponding to the largest eigenvalue F3) is always parallel to the line connecting the arms of the V shape, one minor axis coincides with the symmetry axis, and another is perpendicular to the plane of the V shape. For definiteness, we choose the inplane minor axis (along the symmetry axis) along the line bisecting the central angle formed by the arms of the V shape and pointing away from the vertex, as shown in Fig. 2. It should be noted, however, that our convention of fixing the body frame in such a way that F1 \u2264 F2 \u2264 F3 can result in the interchange of minor axes, e1 \u2194 e2 (and e3 \u2194 \u2212e3 to keep the frame right-handed), upon varying the V-shape opening angle or the aspect ratio h:w as shown in Fig. 2. Considering for definiteness the case of a chubby V shape, as shown in Fig. 2(b), where the only nonzero elements of G are G13 = G31, then Eq. (13) reduces to Uz \u03c9 = C\u0303h s\u03c8s2\u03b8 , (14) where C\u0303h = Ch13 = Ch31 = G13(F1 \u22121 + F3 \u22121)/2 is the pseudochiral coefficient [24,25]. When the same V-shaped propeller is not turning in-sync with the field, the propulsion velocity is given by (see the Appendix) Uz = C\u0303hs\u03c8s2\u03b8 \u03d5\u0307 + G13 ( 1 F3 s\u03c8s\u03b8 \u03c8\u0307 + 1 F1 c\u03c8c\u03b8 \u03b8\u0307 ) . (15) Clearly, for in-sync actuation \u03b8\u0307 = \u03c8\u0307 = 0, \u03d5\u0307 = \u03c9 and Eq. (15) reduces to Eq. (14). When the minor axes interchange [see Fig. 2(a)], similar equations apply, with \u03c8 \u2192 \u03c0/2 \u2212 \u03c8 , \u03b8 \u2192 \u03c0 \u2212 \u03b8 , and G23 replacing G13 everywhere. V-SHAPED STRUCTURES We apply a particle-based method [26] for computing mobility tensors F , G, the resulting anisotropy parameters p , \u03b5, and the chirality matrix Ch for planar V-shaped propellers. This technique is based on multipole expansion of the Lamb\u2019s spherical harmonic solution of the Stokes equations. The object is approximated by (i) a monolayer of N touching rigid spheres of radius a, as shown in Figs", " For straight stripes (\u03b3 = \u03c0 ), the rotation-translation coupling vanishes, G = 0, as can be anticipated from symmetry. Under the convention for the body-frame selection (see Sec. II), for a generic value of \u03b3 , the coupling matrix G has exactly two nontrivial off-diagonal elements 014025-4 Gi3 = G3i, with either i = 1 or, respectively, i = 2, depending on the geometry of the structure. We find G13 < 0 for V shapes with the opening angle \u03b3 = \u03c0/2, while for V shapes with the opening angle \u03b3 = 2\u03c0/3 upon increasing the slenderness, the minor axes interchange, e2 \u2194 e1 (as illustrated in Fig. 2), leading to a sudden change of sign of Gi3 and C\u0303h in Table I, such that G23 > 0 becomes the only nontrivial entry. As one may expect for Stokes flows, the computed rotational and coupling mobilities of hollow 2D structures [in Figs. 3(d) and 3(f)] are quite close to these found for the respective densely packed structures [in Figs. 3(c) and 3(e)]. IV. ACTUATION OF V-SHAPED PROPELLERS BY A CONICAL MAGNETIC FIELD In this section, we consider a number of analytically tractable cases and approximate solutions. In the analysis below, we assume, for definiteness, the orientation of the principal axes as shown in Fig. 2(b). Modification of the present analysis for the slim structures [as in Fig. 2(a)] is straightforward. Assuming magnetization along the symmetry axis e2, i.e., n1 = n3 = 0 and n2 = n\u22a5 = 1, we have = \u03b1 = \u03c0/2. Then from Eq. (11) it follows that \u03c8 = \u00b1\u03c0/2 (it can be readily seen that sin \u03b8 = 0 is not a solution). Substituting these values of the magnetization angles and \u03c8 into Eqs. (10) and (12), we obtain s2\u03b8 = \u00b1 2\u03b4 \u03c9\u0303 ( 1 + \u03b5 \u2212 p\u22121 ) , (16) which imposes a restriction on the value of \u03b4 (or \u03c9) for which an in-sync solution materializes. The \u00b1 signs correspond to two rotational solutions with acute and obtuse wobbling angle \u03b8 , respectively", " This work was supported in part by the Israel Science Foundation (ISF) via the grant No. 1744/17 (A.M.L.) The authors wish to thank Johannes Sachs and Peer Fischer for fruitful discussions. V-SHAPED PROPELLER In the laboratory frame, the translational velocity of a propeller is Ul = RT \u00b7 Ub, where RT is the transposed rotation matrix. From Eqs. (2), we have Ub = G \u00b7 Lb = G \u00b7 F\u22121 \u00b7 b, where the components of the angular velocity b in the body frame are determined by Eq. (5). For the chubby V-shaped propeller [as in Fig. 2(b)], the only nontrivial entries of the coupling matrix G are G13 \u2261 G31. Therefore, the linear velocity in the body frame reads Ub = G13 F3 b 3e1 + G31 F1 b 1e3 . (A1) Thus, the components of the translational velocity in the laboratory frame read Ux = G13 F3 (c\u03d5c\u03c8 \u2212 s\u03d5s\u03c8c\u03b8 ) b 3 + G31 F1 s\u03d5s\u03b8 b 1 , Uy = G13 F3 (s\u03d5c\u03c8 + c\u03d5s\u03c8c\u03b8 ) b 3 \u2212 G31 F1 c\u03d5s\u03b8 b 1 , Uz = G13 F3 s\u03c8s\u03b8 b 3 + G31 F1 c\u03b8 b 1 . (A2) It is seen that, in the in-sync regime, where \u03d5 = \u03c9t + Const, \u03c8 = Const, and \u03b8 = Const [which by Eq. (9) implies that b i , i = 1, 2, 3, are also constant], the components Ux and Uy oscillate with the field frequency \u03c9 and have zero mean upon averaging over a period T = 2\u03c0/\u03c9", " [28] Nondimensionalization with \u03bd0 complies with the definition of the propulsion efficiency \u03b4\u2217 = |Uz, s-o|/\u03bd0 introduced in Ref. [19] that ranks propellers according to their maximal speed (at the step-out) in body lengths per unit time. [29] Sometimes a different limit cycle solution can be obtained; however, its basin of attraction is rather narrow so, for most initial orientations, the solution converges to the same limit cycle. [30] The condition qr > s follows from the fact that the roots of the cubic equation satisfy (\u0303\u03bb1 + \u03bb\u03032)(\u0303\u03bb1 + \u03bb\u03033)(\u03bb\u03032 + \u03bb\u03033) = s \u2212 qr. [31] For a slim V structure [as in Fig. 2(a)], magnetization along e1 results in propulsion similar to that in Sec. IV A subject to some changes due to the interchange of axes, e2 \u2194 e1. The in-sync propuslion velocity in this case is given by Eq. (14) with \u03c8 \u2192 \u03c0/2 \u2212 \u03c8 , \u03b8 \u2192 \u03c0 \u2212 \u03b8 , and C\u0303h = Ch23, resulting in Uz/\u03c9 = \u2212C\u0303h 2\u03b4/(1 + \u03b5 \u2212 p\u22121). Since C\u0303h changes sign, the propulsion in the +z direction requires magnetization along \u2212e1. [32] J. E. Marsden and T. J. R. Hughes, Mathematical Foundations of Elasticity (Dover, New York, 1983). 014025-1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003301_j.matpr.2020.12.1110-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003301_j.matpr.2020.12.1110-Figure1-1.png", "caption": "Fig. 1. Reference hub model.", "texts": [], "surrounding_texts": [ "The failure of rim wheel is due to fatigue failure. For the improvement of fatigue life, the material optimization and design optimization to find parametric design which gives better fatigue rate. So we experiment and analyze the stress strain and deformation values for different kinds of materials." ] }, { "image_filename": "designv11_22_0002205_lwc.2020.2965445-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002205_lwc.2020.2965445-Figure1-1.png", "caption": "Fig. 1. General 3-D structure excited by an impressed current source J.", "texts": [ " In Section III, the accuracy and effectiveness of the proposed method will be demonstrated by examining the matrix condition and comparing the results with that of the 0018-9464 \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. original double curl equation. Finally, the conclusion is drawn in Section IV. Consider a general 3-D boundary value problem, as shown in Fig. 1. Assume that the structure is inhomogeneously composed of three bodies, 0, 1, and 2, among which 0 is bounded by D (solid line) and N (dashed-dotted line); 1 and 2 are bounded by 1 and 2, respectively. In addition, the structure is excited by an impressed current source J. Thus, A satisfies the following double curl equation with generalized Coulomb gauge: \u2207 \u00d7 1 \u03bc \u2207 \u00d7 A \u2212 \u03b5\u2207 1 \u03b1\u03b52\u03bc \u2207 \u00b7 \u03b5A = J (1) which is the static case of the A-equation in [19], with \u03b5, \u03bc, and \u03b1 being the permittivity, permeability, and gauge factor, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003257_s11665-020-05402-8-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003257_s11665-020-05402-8-Figure4-1.png", "caption": "Fig. 4 Total deformation of redesigned turbine", "texts": [ " The curling of the blade tips increases mechanical strain and results in higher stress concentration at the support structure and at the first layers of the blade tips. The geometry was redesigned to integrate a horizontal disk that transitioned into all the blades to reduce stress, strain and deformation. This modification in the design of the conventional turbine allows heat to transfer and reduce thermal strain that causes higher stress and/or deformation of the blade tips. The AM process simulation was re-performed. Figure 4 shows the redesigned (V2) turbine build at 48.02 min. A comparative plot, Fig. 5, shows deformation data from the conventional turbine (V1) in blue and the redesigned turbine (V2) in orange. By adding a geometric disk feature at the base of the turbine, a significant decrease in deformation was observed, which reduces the internal stress. This turbine was printed via an EOS M100 printer as explained in Sect. 4.5. Journal of Materials Engineering and Performance Volume 30(2) February 2021\u20141179 Recycled scrap material made from Inconel 718 was used to manufacture metal powder alloy (IN718) using a gas atomization process patented by MolyWorks Materials Corporation (Ref 17)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003519_s10846-021-01359-5-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003519_s10846-021-01359-5-Figure4-1.png", "caption": "Fig. 4 Custom designed and manufactured leader robot with key components labeled", "texts": [ " These wheels and rollers are 3D printed using PLA material and are then assembled using Allen bolts. To improve traction between the rollers and the ground surface, heat shrink of polyolefin material is applied to the rollers. Additionally, flange couplings are custom-designed for the wheels and motors, and machined through lathe operations. The upper platform is a turntable mechanism containing thrust bearings and 3D printed mounts. Furthermore, clamping jaws are provided on the turntable to hold the payload intact, as shown in Fig. 4. Figure 5 shows the control architecture of the leader robot. It consists of a Wi-Fi development kit called as Wemos D1 mini running at 80MHz with ESP8266 chip. The leader robot is remotely controlled from a mobile phone using Blynk App, as seen in Fig. 9. The app contains two analog joysticks which control the translation and rotation of the robot, respectively. Wemos, containing a Wi-Fi module also has a few input-output pins. These are used to signal two channel motor drivers powered by a Li-Po batter of 11" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001174_pime_proc_1961_175_037_02-Figure22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001174_pime_proc_1961_175_037_02-Figure22-1.png", "caption": "Fig. 22.-continued", "texts": [ " Types of failure and significant variables The results of the fatigue tests on various bearing assemblies are given in Table 6, which, unless otherwise stated, shows the conditions at which the fatigue failure was first observed. In assessing the significance of these results careful recognition must be made of the five most important variables affecting them, which are: type of metal, thickness of bearing shell, type of connecting-rod, type of oil grooving and oil holes, and bearing clearance. The engine connecting-rod used for the thin shell bearing is that shown in Fig. 22b. It is substantially more rigid than the original assembly, Fig. 22a for thick shell bearings. The two distinct areas of material break up of a typical failure of ungrooved bearing are shown in Fig. 27. On grooved bearings the areas of fatigue were closer to the groove, so that in both types the initial cracks started not at the positions of maximum oil film pressure bur rather at positions of maximum pressure gradient across the bearing. The fatigue cracks seldom extend to the outer edges. The most significant points arising from these results may be summarized as follows : influence of housing rigidity The influence of housing rigidity is demonstrated by comparison of the copper-lead bearings Nos" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000465_0954406219878755-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000465_0954406219878755-Figure4-1.png", "caption": "Figure 4. Coordinate systems for meshing of enveloping conical worm pair.", "texts": [ " In equation (15), the notation R\u00bd ~k1, \u2019 is the rotation transformation matrix.15 Meshing of enveloping conical worm pair The working meshing of an enveloping conical worm pair is completely identical to the cutting meshing of the conical worm wheel, whereupon these two concepts will not be differentiated in the following. During machining a conical worm wheel, a stationary coordinate system o2 O2; ~io2, ~jo2, ~ko2 is utilized to denote the initial location of the conical worm gear as described in Figure 4. The unit vector ~ko2 coincides with the axis of the worm gear, and its positive direction is from the little end to the big end of the worm wheel. On the other hand, the unit vector ~ko1 is perpendicular to the unit vector ~ko2 because this paper only considers the orthogonal enveloping conical worm pair. The unit vector ~io2 is along the common perpendicular of the two vectors ~ko1 and ~ko2. The length of O2O 0 ! is the shortest distance between the two unit vectors ~ko1 and ~ko2 along their common perpendicular line and O2O 0 ", " Similarly, in o1, the normal vector ~n of S\u00f0 \u00de 1 can be obtained from equation (6) by dint of the coordinate transformation, and the result is ~n o1 \u00bc R ~ko1, \u20191 \u2019 h i ~n o1 \u00f017\u00de Without loss of generality, it is also possible to presume that the angular speed of the worm rotation about ~ko1 equals to 1 rad=s. According to the meshing theory of gearing,15 in o1, it is facile to acquire the speed vector of relative movement between the worm and worm wheel as follows ~V12 o1 \u00bc Vx ~io1 \u00fe Vy ~jo1 \u00fe Vz ~ko1 \u00f018\u00de where Vx \u00bc y o1 1 i12 zow \u00fe z0\u00f0 \u00de, Vy \u00bc x o1, and Vz \u00bc x o1 \u00fea i12 . Inside z0 \u00bc zA \u00fe Lw 2 . Herein the two notations Lw and zA express the thread length of the worm and its installing distance along its center axis as shown in Figure 4, respectively, and zA \u00bc kAa, in which kA is a coefficient. According to the theory of gearing,16 the meshing function, , for the enveloping conical worm drive can be acquired from equations (17) and (18). By feat of the vector rotation technique to simplify the result achieved above,15 the obtained final outcome is \u00bc ~n o1 ~V12 o1 \u00bc 1 i12 A sin \u20191 \u2019\u00f0 \u00de\u00bd \u00fe B cos \u20191 \u2019\u00f0 \u00de \u00fe C \u00f019\u00de wherein A\u00bc ny zow\u00fe z0\u00f0 \u00de nzyow, B\u00bc nx zow\u00fe z0\u00f0 \u00de \u00fe nzxo1, C\u00bc i12nyp sin 1\u00fe nz a i12p cos 1\u00f0 \u00de \u00f020\u00de Through the coordinate transformation from o1 to 2, the equation for the worm wheel tooth face, S\u00f0 \u00de 2 , (S \u00bc 1, 2) can be attained from equations (16) and (19), and expressed in 2 as ~r2 2 \u00bcR ~k2, \u20191 i12 O2O1 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000237_s00216-019-01884-1-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000237_s00216-019-01884-1-Figure3-1.png", "caption": "Fig. 3 The typical FCS strategy of protein detection", "texts": [ " applied FCS to investigate the average rate constants of single base-pair hybridization/dehybridization of a dsDNA at different temperatures and proposed the stretched exponential zipper (SEZ) model [93]. Hu et al. developed a sensitive and microscale method to study the formation and stability characterization of triplex DNA using FCS. The principle is mainly based on the excellent capacity of FCS for sensitively distinguishing between free ssDNA fluorescent probes and fluorescent probe dsDNA-hybridized complexes [94]. Quantitative analysis on proteins is essential in the research of life science. The principle of homogeneous analysis on proteins is similar to that on DNA as shown in Fig. 3. A fluorescent-labeled probe that can specifically bind with the target protein was used. The fluorescent probes appear a short characteristic diffusion time in the detection volume before binding. Once binding, a complex component with bigger characteristic diffusion time emerges. The determined average diffusion times from auto FCS measurements increase with more probes bound with target molecules. The established standard linearship between diffusion times and concentration of target molecules can be used for its quantification" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002701_aero47225.2020.9172614-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002701_aero47225.2020.9172614-Figure1-1.png", "caption": "Figure 1: Models used for the prototyping of the Tetracopter elementary module.", "texts": [ " A single rotorcraft is called an elementary module and several elementary modules can assemble at different scales to form a Fractal Tetrahedron Assembly (FTA) based on the Sierpinski tetrahedron [4]. The goal of the FTA design is to provide rigidity in three- dimensions while maintaining the efficiency of a virtually two-dimensional modular robotic flight system. Different designs of the elementary module can be considered. In this work, we consider an elementary module consisting of a quad-rotorcraft with a regular tetrahedron structure, as seen in Fig. 1a. In practice, this tetrahedral quad-rotorcraft, which we name Tetracopter, is made of four identical submodules and a common payload including the battery and electronics of the quad-rotorcraft. Fig. 1b shows this submodule, which has only one rotor and is therefore uncontrollable by itself. The results of this research are the foundation of future research of Georgia Tech\u2019s Decision and Control Laboratory that aims to develop tetrahedral rotorcraft that can self-assemble in flight. This paper is organized as follows: \u2022 The details of the Fractal Tetrahedron Assembly are presented in Section 2, along with the analyses and results that justify this assembly scheme. Among them are a study of the dynamics of the assembly and some simulated results of the internal forces of the assembly structure in different situations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001169_icra.2014.6907355-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001169_icra.2014.6907355-Figure3-1.png", "caption": "Fig. 3. A quad-arm manipulator holding a single object", "texts": [ " The center of the grasped object is always controlled by hybrid position/force control law in order to keep the previously specified internal force between the two robots while moving. In this research, the extended hybrid position/force controller is derived based on the kinematics and the statics of the four arms. In this section, the kinematic and static equations are derived using the task vectors proposed in [2], which involves generalized forces, velocities and positions. Let us consider four arms holding an object as shown in Fig. 3. \u2211o, \u2211hi and \u2211a represent the base frame, hand frame of arm i, and the object frame, respectively. The object frame is fixed to the object. The virtual sticks olhi (i = 1 \u223c 4) are determined at the moment the robots hold the object as vectors from Oi to Oa. The virtual sticks is fixed to the hand frame \u2211hi. Let h1 and h2 be two arms of one robot, while h3 and h4 be two arms of the other one. The deformation of the object and the slippage of the hands on object are assumed to be very small in this research, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002546_ilt-04-2020-0143-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002546_ilt-04-2020-0143-Figure1-1.png", "caption": "Figure 1 3D-pin samples with (a) solid structure, (b) internal square structure, (c) internal rectangular structure, (d) internal circular structure, (e) internal triangular structure and (f) internal triangular flip structure", "texts": [ " Therefore, although surface roughness plays an important role in reducing COF over time, it does not have a significant effect on the tribo-mechanical properties of 3D-printed ABS with different internal geometries. The printing parameters were selected based on the optimum values of the previous study (Norani et al., 2020). Since the printing process parameters for all samples are similar in this study and thus surface roughness is assumed to be constant within a specific range and can be neglected. From Figure 1, the sample was categorized by six F1 different internal geometries. The dry sliding test was performed using a pin-on-disc tribometer. The pin sample was slide against carbon chromium steel disc according to the ASTMG99-0, 2016. The test was run at a constant applied load 58.68N, sliding speed 600 rpm, sliding distance 1000m. All tests were performed at 27\u00b0C. Before the test, acetone was used to clean the pin and disc. The COF and wear datawere collected directly from the experiment. The compression test was performed using the Instron compression test machine at a rate of 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000655_j.prostr.2019.12.049-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000655_j.prostr.2019.12.049-Figure4-1.png", "caption": "Figure 4. FEA model of AM In718 fatigue test sample shows uniform stress generated across the gauge region. Strain values were measured and confirmed to be within 3% of the predicted value.", "texts": [ " The tests were focused on demonstrating fatigue enhancement of In718 after 350 hours exposure at 600oC regarding jet engine turbine applications. In general, there is concern for AM samples about voids and inclusions in the materials that might be randomly distributed and thereby initiate failure over a large region [16]\u2013[19]. The 4-point bend testing configuration is especially effective because it provides a very uniform test stress over a large fraction of the gauge section and thus samples relatively large volumes of the material. Figure 4 shows our finite element analysis prediction of the stress and strain developed during loading of the test samples. As can be seen, the test setup develops very uniform strain over the entire region between the load rollers. to begin the tests, strain gauges were applied to a first sample, the sample loaded on the machine and a 7000 MPa (1000 lb) load applied. For this load the strain gauge showed 615 microstrain which was within 3% of the 595 microstrain predicted indicating that the applied stress was accurately calibrated to the loading and was consistent with our modeling. In operation stress loadings were reported to this measured/predicted load-to-stress calibration result. The analysis shown in Figure 4 gives a maximum stress along the gauge section of 134 MPa (19.170 ksi) for an applied load of 4.448 kN (1000 lbs). For the fatigue test results shown in this report a maximum load of 22.65 kN was applied(5093 lbs) which generated a maximum stress of 683 MPa (97.6 ksi). Literature values for the yield stress of AM In718 [20] give a value of 791 MPa (113 ksi) at 0.2% yield. Literature values from Special Metals Corporation [21] give a yield stress of 1050 MPa (150 ksi) at 0.2% elongation rating the AM material at 24% lower yield strength. Figure 4. FEA model of AM In718 fatigue test sample shows uniform stress generated across the gauge region. Strain values were measured and confirmed to be within 3% of the predicted value. With the basic samples fabricated, specific surface treatments were applied including single shot peening (0.3 mm size shot with medium A-scale intensity), dual shot peening with a light and then heavy peening (0.3 mm size shot with medium A-scale intensity followed by peening with 0.1 mm shot size with high C-scale intensity), and standard laser peening", " AM In718 fatigue test samples manufactured for 4-point bend testing. The 4-point bend test configuration provides uniform stress across a wide range of the gauge section. In general, there is concern for AM samples about voids and inclusions in the materials that might be randomly distributed and thereby initiate failure over a large region [16]\u2013[19]. The 4-point bend testing configuration is especially effective because it provides a very uniform test stress over a large fraction of the gauge section and thus samples relatively large volumes of the material. Figure 4 shows our finite element analysis prediction of the stress and strain developed during loading of the test samples. As can be seen, the test setup develops very uniform strain over the entire region between the load rollers. to begin the tests, strain gauges were applied to a first sample, the sample loaded on the machine and a 7000 MPa (1000 lb) load applied. For this load the strain gauge showed 615 microstrain which was within 3% of the 595 microstrain predicted indicating that the applied stress was accurately calibrated to the loading and was consistent with our modeling. In operation stress loadings were reported to this measured/predicted load-to-stress calibration result. The analysis shown in Figure 4 gives a maximum stress along the gauge section of 134 MPa (19.170 ksi) for an applied load of 4.448 kN (1000 lbs). For the fatigue test results shown in this report a maximum load of 22.65 kN was applied(5093 lbs) which generated a maximum stress of 683 MPa (97.6 ksi). Literature values for the yield stress of AM In718 [20] give a value of 791 MPa (113 ksi) at 0.2% yield. Literature values from Special Metals Corporation [21] give a yield stress of 1050 MPa (150 ksi) at 0.2% elongation rating the AM material at 24% lower yield strength" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001601_s00170-016-8487-6-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001601_s00170-016-8487-6-Figure2-1.png", "caption": "Fig. 2 Consecutive steps of a typical Exp-ECAE operation", "texts": [ " 1 illustrates, an Exp-ECAE die simply involves a couple of perpendicular cylindrical channels which intersect at a spherical cavity. The centerlines of channels are located next to the center of spherical hollow with a distance of e=2 mm (Fig. 1). At the beginning of process, the punch pushes the billet with an initial diameter of D0 until the billet reaches the ballshaped region. Then, the bar gradually expands and fills the hollow. Afterwards, the metal flows from the cavity to the exit channel and is, therefore, extruded to its initial diameter [13]. When the billet is placed in the inlet channel (Fig. 2a), by the punch movement, the specimen reaches the spherical hollow and, due to the eccentricity of e, touches the die wall at a single point. This causes some bending in the billet (Fig. 2b). The globe-shaped region prevents the material from progressing toward the exit channel then the billet has to expand and fill the cavity (Fig. 2c). After filling the spherical part of tool, the metal is extruded toward the exit channel and recovers its initial diameter (Fig. 2d). The required pressure for this step of the Exp-ECAE acts as a back pressure for the expansion of the remaining material in the inlet channel. Therefore, the material follows the described sequences throughout the whole process. Further passes can be conducted by rotating the die by 90\u00b0 to exchange the positions of the vertical and horizontal channels then running the punch into the new vertical channel. Exp-ECAE imposes plastic strain into the sample during the expansion, the extrusion, and the shear stages" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003664_tia.2021.3084549-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003664_tia.2021.3084549-Figure5-1.png", "caption": "Fig. 5. Structure of a consequent pole type rotor (Field weakening).", "texts": [ " An SMC is excited in a direction opposite to the PM magnetization direction by field winding and then forms a magnetic pole of the opposite polarity to that of the PM. We refer to the direction of the field winding current as the field strengthening direction. By contrast, the direction of the field winding current excited in the same direction as the PM is referred to as the field weakening direction. Fig. 4 shows the flow of the field magnetic flux in the field strengthening direction. It is used in a wide operating area above a certain level of torque. The higher the torque that is required, the larger the field current that is supplied. Fig. 5 shows the flow of the field magnetic flux in the field weakening direction. It is used at TABLE I. SPECIFICATION OF THE PROPOSED MOTOR Number of pole 20 Number of slot 24 Outer diameter (including coil end) 250 mm Axial length 76.4 mm Volume (including coil end) 3.75 L Air gap 1.0 mm PM thickness 13.0 mm Pole arc angle of PM 21 degrees Pole arc angle of SMC 15 degrees Turns per coil of armature winding 9 turns Maximum armature current 74.77 Arms Maximum armature current density 11.90 A/mm2 Turns per coil of field winding 280 turns Maximum field current 5" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003543_j.optlastec.2021.107153-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003543_j.optlastec.2021.107153-Figure3-1.png", "caption": "Fig. 3. Schematic diagram of DLBSW technology.", "texts": [ " 2, The laser beam is produced by a 12,000 W disc laser and homogeneously transmitted to two laser welding head through two independent optical fibers, respectively. During the welding process, the weld beads of the Tjoints were made simultaneously from both sides of the stringer. Hence, the laser beam and shielding gas were delivered symmetrically along the center-line of the stringer. Furthermore, the offset angle between laser head and horizontal plane was set to 40\u25e6 in order to stabilize the doublesided welding process. The schematic diagram of DLBSW is visible in Fig. 3. According to the reference, Ti6Al4V alloy is easily oxidized to form dense oxide film when exposing to air [22]. Therefore, the stringers and skins were polished carefully before welding process. Additionally, the oil impurities and metallic dust on the surface of the panels were cleaned thoroughly with mechanical and chemical cleaning methods. There were three well-formed skin-stringer constructions chosen for the sample preparation, and their specific welding parameters are given in Table 3. After welding process, the metallographic specimens were cut along the cross-section of T-joints, and the uneven area shall be avoided" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000607_j.rcim.2019.101913-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000607_j.rcim.2019.101913-Figure4-1.png", "caption": "Fig. 4. Quadratic B\u00e9zier Curve.", "texts": [ " The edge weights consequently are not the Euclidean distance between two nodes, but the arc length of each B\u00e9zier curve, given by \u222b \u239c \u239f= \u239b \u239d \u239e \u23a0 + \u239b \u239d \u239e \u23a0 D d dx B t d dy B t dt( ) ( ) , 0 1 2 2 (2) where D is the arc length of each parametric curve. Besides, each maneuver requires handling and pushing stages. In other words, the robot has to travel towards the object and touch it, then guarantee it will not miss robot-box contact during their displacement towards the destination. Mathematically, two B\u00e9zier curves are required. The first one indicates the robot-object path, where the start node p0 is the robot position, the auxiliary point p1 is in some place of an orthogonal line that intersects the opposite box surface (see Fig. 4), and finally, p2 is the midpoint at the opposite box surface. Further, the second curve starts at the last point of the first curve, it has the mirrored p1 as its auxiliary point, and it ends at the goal point. Once completing the graph representation, we should find the shortest path to handle and push each object to its destination. Thus, Dijkstra\u2019s algorithm calculates the best route taking into account the robot\u2019s home position, the goal localization, and the position of all objects as they are \u201cseen\u201d by the robot" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001517_iet-cta.2014.1144-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001517_iet-cta.2014.1144-Figure2-1.png", "caption": "Fig. 2 Generalised configuration coordinate a Simple pendulum", "texts": [ " Fortunately, its stability can be verified for many practical non-linear systems, such as pendulum, n-link robotic arm and autonomous underwater vehicles. In future, a non-linear version of Z, that is, Z(x\u0304) is planned to be investigated, to deal with non-linearities f ( \u02d9\u0304x, x\u0304) in general. In the following subsections, we apply our proposed L2 neuroadaptive controller to some benchmark dynamic systems. Unlike previous works, the major difference in the examples considered here is that the uncertainties are no longer in LIP form. In this example, as shown in Fig. 2a, \u03b8 is the generalised configuration coordinate, that is, q = \u03b8 and G = 1. Assume that there is no friction, that is, D = 0. u is the input torque. Further description of the system is given as M = ml2 (71) The Hamiltonian of the system is given by H (\u03b8 , p) = KE + PE = 1 2 p2 ml2 + mgl(1 \u2212 cos \u03b8) (72) where p = ml2\u03b8\u0307 . The nominal magnitudes of pendulum parameters and uncertainties are given in Table 1. The objective is to stabilise the system at \u03b8d = (\u03c0/2) rad. As a straightforward choice, (\u03b8 , t) = \u03b8\u0304 = \u03b8 \u2212 \u03b8d ", " This controller performance is achieved within the PCH formalism only by incorporating the IP filtering and unlike [12], no extra transformation is needed. Furthermore steady-state error is minimised without any integral action in the control law. IET Control Theory Appl., 2015, Vol. 9, Iss. 12, pp. 1781\u20131790 \u00a9 The Institution of Engineering and Technology 2015 A fully actuated two-link, two degrees of freedom planner manipulator is a benchmark mechanical system . This robot arm has all the non-linear effects common to general robot system. Consider a manipulator, shown in Fig. 2b, with masses mi and link lengths ai for i = 1, 2. The generalised configuration coordinates are link angles \u03b81 and \u03b82, and therefore q = \u03b8 = [\u03b81 \u03b82]T. The system is described by (55) with u = [ \u03c41 \u03c42 ]T , G = I2, D = diag(d1, d2) M (\u03b8) = [ M11(\u03b8) M12(\u03b8) M21(\u03b8) M22(\u03b8) ] where M (\u03b8)11 = (m1 + m2)a 2 1 + m2a2 2 + 2m2a1a2 cos \u03b82 M (\u03b8)12 = M (\u03b8)21 = m2a2 2 + m2a1a2 cos \u03b82 M (\u03b8)22 = m2a2 2 Note that \u03b8\u0304 = \u03b8 \u2212 \u03b8d . The manipulator works in the horizontal plane, therefore gravity effects are negligible and V (\u03b8) = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000353_iemdc.2019.8785381-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000353_iemdc.2019.8785381-Figure6-1.png", "caption": "Fig. 6. A sketch of the space vector diagram during the rotor rotation with constant speed, stator voltage and frequency.", "texts": [ " This method is meant to keep constant dq stator voltages during the rotation, in this way abc voltages will have a perfectly sinusoidal behavior in time. In the other hand, the stator dq current will change in every step, leading to time harmonics in abc currents. If the motor is skewed, the multislice skewing analysis is necessary in each time step. Actually, in the sequence of steps above, the inductances are computed neglecting the rotor skewing, since the variation is not serious. The skewing analysis is made after the currents computation, in order to have a better estimation of the torque, saving computational time. Fig. 6 shows one time instant of the motor space vectors, during steady-state under-load operation. The graphic points out that the stator current and flux linkage have an average value overlapped to high frequency components along the RFO reference frame d and q. These harmonic components are due to the slotting and to the saturation related to stator space harmonics. As regards the rotor winding, current and flux linkage space vectors can pulsate only along q and d direction respectively. This derives from the rotor current computation method itself: in each rotor position, the rotor q-axis current is derived to lead to zero \u03bbrq, whereas ird is forced to be zero" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002798_s00170-020-06152-6-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002798_s00170-020-06152-6-Figure13-1.png", "caption": "Fig. 13 Different building directions for single part", "texts": [ " Geometrical tolerancing and datum establishment in design are strongly related with part\u2019s functionality as well as its assembly, especially in cases of multi-component assemblies. In most cases there is a strong connection between manufacturing process (plan), geometrical tolerancing, and especially datum establishment. Furthermore, in the case that in one single AM build, multiple parts are produced, not only laterally but vertically too, all these factors may affect differently the produced parts (Fig. 13). Last but not least, geometric and surface quality of AM parts are affected by the build direction due to the stepwise discretization of layer thickness [31\u201333]. Most common defects of popular AM technologies (more usual for FDM, SLA and DLP) are shrinkage, warpage, and adhesion defects. Although most manufacturers of 3D printing equipment constantly develop techniques to compensate them, all these types of defects lead to poor surface quality making it difficult to inspect and evaluate relevant GPS" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000828_j.amc.2015.05.111-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000828_j.amc.2015.05.111-Figure1-1.png", "caption": "Fig. 1. Control mechanism of main rotor.", "texts": [ " When the y-coordinate of any wheel center in the inertial frame (seen in 2.2) is less than zero, it means the wheel has already touched the ground and the corresponding damping and spring coefficients are nonzero. Otherwise, when the helicopter is flying in the air, each coefficient can be set to zero and the airframe (except the main and tail rotors) is regarded as a rigid body with mass symmetry about the central plane. (2) The helicopter in this paper has an articulated rotor which is shown in Fig. 1. The main rotor has four ideal-twist blades whose angular velocity and chord length are all constants. The tip loss and lead-lag motions of blades have been ignored. (3) The tail rotor thrust which balances the antitorque from main rotor can be adjusted by the model automatically and the antitorque from tail rotor has been ignored. In addition, the aerodynamic drag coefficient is considered a constant. 2.2. Interpretation of coordinate system (1) The inertial frame OXY Z: this coordinate system is attached to the earth, which is coincident with body-fixed frame at initial time, its original point is 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001279_s0263574714000149-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001279_s0263574714000149-Figure4-1.png", "caption": "Fig. 4. The kinematic chain model. (a) A kinematic chain of bodies (KCB). (b) Mobility of a mechanism with closed loops after opening and replacing two joints with equivalent constraints.", "texts": [ " 3), which enables an easy identification of loops. This model9 is very popular, especially in the kinematic and dynamic analysis of robots and other controlled mechanical systems as it facilitates the direct control of joint motion parameters. Within this model, the KCB is thus an open chain of bodies serially interconnected through joints and represents the structural primitive building block of a mechanism. Its dof is variable with the number and type of individual joints integrated by the KCB (Fig. 4(a)). This is in contrast to the http://journals.cambridge.org Downloaded: 18 Mar 2015 IP address: 128.233.210.97 MBS structural model where the primitive entity (the body) has a constant number of dof\u2019s (S = 6 in space and S = 3 in plane). For this reason, after defining the KCB structural model of a mechanism, a first step towards calculating the mobility is to determine the number of degrees of freedom for each individual KCB. This is given by the dof of the terminal body that cumulates the degrees of freedom of the preceding joints (Fig. 4(a)). The KCB dof formula is thus: DOF = \u2211 fi. (5) The mobility of the mechanism is then calculated in two steps; first, the mechanism is discretized into individual KCBs by opening a joint in each closed loop (Fig. 4(b)); secondly, the mobility is calculated similarly with the MBS model\u2014as a sum of KCBs degrees of freedom that is diminished with the number of kinematic constraints between the various KCBs. In this way the mobility relationship for the closed loop mechanisms is: M = \u2211 fi \u2212 \u2211 rk (6) in which \u2211 ri is the number of constraints corresponding to the opened joints. Equation (6) is a mobility formula previously proposed by various authors, such as Freudentstein and Alizade,10 who also made the connection with the kinematic analysis" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002623_acc45564.2020.9147323-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002623_acc45564.2020.9147323-Figure3-1.png", "caption": "Fig. 3. Illustration of the soft robot and degree of curvature estimation using motion capture data of the markers.", "texts": [ " First the experimental setup is described. Subsequently, the experimental results are presented and discussed. The experimental setup shown in Fig.2 consists of a pneumatically actuated pleated type soft robot [24], OptiTrack motion capture system and two i7 16GM RAM Windows 10 laptops: one running the control algorithms on Matlab 2015 and the other simulating the rigid master robot. Experiments with a physical rigid master robot will be reported in the future. 1) The Soft Robot: The soft robot used in the experiments, as shown in Fig.3(a), comprises of three bi-directional pleated type segments (n = 3) which were fabricated following methods outlined in [24]. The robot was constrained to move on a horizontal table and ball transfers were used underneath near the segment joints to reduce friction. Each segment has two compartments that are individually actuated pneumatically. The segments are assumed to deform with a constant strain under the applied pressure, thus, having a constant curvature. The middle layer of each segment (the joint between the two chambers) is inextensible due to the restrained material layer", " The segment lengths along the inextensible middle 2126 Authorized licensed use limited to: Cornell University Library. Downloaded on September 11,2020 at 07:03:17 UTC from IEEE Xplore. Restrictions apply. layer were measured to be Lsi = 0.125 m. The segment masses msi = 0.110 kg were measured prior to joining the segments together. The material properties of each segment were assumed to be identical. Therefore, for all the segments identical torsional stiffness of ki = k and damping of di = d were assumed. The identification of the nominal values of k and d will be discussed subsequently. As seen in Fig.3(b), an unactuated end effector resembling a soft gripper of mass me = 0.050 kg was attached to the tip of the soft robot. 2) Degree of Curvature Estimation using Motion Capture: The positions of the end points of each segment was extracted using a Motion Capture system (OptiTrack) by attaching clusters of markers on the segment ends as shown in Fig.3(b). These labeled markers were assumed to be lying on the horizontal plane throughout the trial. The base of the soft robot was securely attached to the table so that experiments could be conducted without re-calibration. The degree of curvature of each segment was calculated using the properties of the dot product of the labeled marker positions as illustrated in Fig.3(c): qsi = 2 ( \u03c0 2 \u2212 cos\u22121 ( (a\u2212b)\u00b7(c\u2212b) ||a\u2212b|| ||c\u2212b|| )) . 3) The Actuation Unit: The soft robot was actuated using a pneumatic controller unit based on the open source hardware platform [25] with command inputs serially transmitted to the control board. The compressed air to the unit was supplied by an external compressor at a constant pressure of 20psi. The air pressure in the segments was regulated by Pulse-Width Modulation (PWM) using a frequency of 100Hz. The control input calculated by the controller, in terms of a torque, was converted to a PWM signal for each segment using corresponding mappings" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002384_s12206-020-0325-y-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002384_s12206-020-0325-y-Figure1-1.png", "caption": "Fig. 1. The hybrid helical gear fabricated with steel and aramid/phenol composite materials.", "texts": [ " For the fabrication of a hybrid gear, phenol resin was infiltrated into the dry preform sheet located between the inner and outer steel parts in advance, by pressing a hot press where temperature is controllable. The stacked dry hollow preforms impregnated with phenol resin (i.e., aramid/phenol composite) is then cured insitu. During the cure, the phenol resin with chopped aramid fibers from preforms can flow and be inserted into complicated interface contours of steel parts. Several hybrid gears were successfully produced by the wet laid manufacturing with a specially designed hot-pressing mold and one hybrid gear is shown in Fig. 1. The gear dimension is as follow: r1 (steel shaft inner radius) = 18.5 mm, r2 (steel shaft outer radius) = 21.5 mm, r3 (composite outer radius) = 34.4 mm, r4 (dedendum radius) = 36.5 mm, r5 (addendum radius) = 41.05 mm, and the width of a gear, t = 13.2 mm, as shown in Fig. 1. To secure the interface adhesion between steel and aramid/phenol composite, the shapes of interface are complicated and in bulgy and hollow contour, as shown in Fig. 2. A torsion test was performed for the hybrid gear with an initial interface feature in Fig. 5 in order to determine the failure torque and to investigate the interface failure mode between the composite and steel parts. Torque increases gradually through the central shaft of a hybrid gear until torque failure occurs while the teeth region is fixed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.5-1.png", "caption": "Fig. 3.5 Targeting with one degree-of-freedom (DOF)", "texts": [ "9) Note: In mathematics and physics, the square root of \u22121 is typically denoted with i , whereas in many technical areas j is used. In both Python and Matlab, j can be used: x = -1+0j np.sqrt(x) >>> 1j In polar coordinates, the similarity between one-dimensional translations and singleaxis rotations becomes obvious (Fig. 3.1). Task: If a gun originally pointing straight ahead along the +x axis is to shoot at a target at P = (x, y), by which amount does the gun have to rotate to point at that target (Fig. 3.5)? Solution: The gun barrel originally points straight ahead, so the direction of the bullet aligns with nx. The rotation of the gun is described by the rotationmatrixR = [ cos \u03b8 \u2212 sin \u03b8 sin \u03b8 cos \u03b8 ] =[ n\u2032 x n \u2032 y ] . The direction of the gun barrel after the rotation is given by n\u2032 x = p |p| , which is also the first column of the rotation matrix R. 34 3 Rotation Matrices Combining these two equations leads to the vector equation ( cos \u03b8 sin \u03b8 ) = 1\u221a x2 + y2 ( x y ) . They-component of this vector equation is sin \u03b8 = y\u221a x2+y2 \u2192 \u03b8 = arcsin y\u221a x2+y2 Note: For small angles (\u03b8 1), sin(\u03b8) and cos(\u03b8) can be expanded with a Taylor series, and one obtains in a linear approximation sin(\u03b8) \u2248 \u03b8 + 0(\u03b82), and (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001651_s00170-016-8545-0-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001651_s00170-016-8545-0-Figure1-1.png", "caption": "Fig. 1 Manufacturing processes of shrouded blisk", "texts": [ " There are several factors that influence the geometric accuracy of EBM most such as the diameter of metal powder. And an EBM machining accuracy of \u00b10.3 mm has been reported [17]. A combined method of EBM and EDM for shrouded blisk manufacturing is presented in this paper, in which EBM is used for generating the near-net forming blank of blisk and EDM is responsible for finishing. This paper will concentrate on the manufacturing process of the EBM/EDM combined method and other related technical matters. Figure 1 shows the machining processes of the regular forging/EDM machining method and the EBM/EDM combined method, respectively. In both blisk machining processes, EDM just takes the machining work of the blisk blades. Other profiles except the blades are mainly finished by cutting. Generally, blade profile must achieve very high shape and position accuracy after the EDM process. The main difference is that the EBM blank has rough profile of the blades on it. As shaped electrode is used, the profile of the workpiece is copied from the electrode after the unwanted material is eroded by discharges" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003519_s10846-021-01359-5-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003519_s10846-021-01359-5-Figure7-1.png", "caption": "Fig. 7 Hardware assembly for sensing a displacement based force from leader robot. Liner potentiometer is used to sense d whereas rotary potentiometer senses \u03b8", "texts": [ " In the front acrylic sheet, there are 3D printed mounts for ultrasonic sensors inclined at an angle of 60\u25e6, as shown in Fig. 6. The upper turntable has an additional constrained degree of freedom in a linear direction towards the leader robot\u2019s force. Furthermore, the turntable is coupled with a pair of spur gears in 1:2 ratio to sense the angle at which the RL applies force. Hence, both the magnitude (d) and direction of the force (\u03b8 ) are sensed in this manner, which can be] seen in the assembly displayed in Fig. 7. Figure 8 displays the overall control structure of the follower robot. The ultrasonic sensor is mounted in an inclined fashion so to detect obstacles in the forward-right and forwardleft positions of the robot. Additionally, the upper platform contains a 40mm B50K\u03a9 slide potentiometer, arranged such that the slide is along the direction of force of the leader robot. The turntable is coupled with a 10K\u03a9 rotary potentiometer through a pair of spur gears, as shown in Fig. 6. All these are sensed by an Arduino UNO running at 16MHz" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001507_iccas.2015.7364708-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001507_iccas.2015.7364708-Figure5-1.png", "caption": "Fig. 5: The estimated frequency of KCSF-20: radial direction (top) and axial direction (bottom)", "texts": [ "23 times greater than that of the CSF-20. Fig. 4 depicts the natural frequencies of two flex splines measured along the radial direction. The frequency ratio between the KCSF-20 and CSF-20 measured along the radial direction was the same with the result along the axial direction. The natural frequencies of KCSF-20 were obtained in axial and radial directions by Finite Element Model (FEM). In order to obtain the FEM, the flex spline was 978-89-93215-09-0115/$31.00@ICROS modelled with C3D8I solid elements. Fig. 5 shows the FEM of the flex spline and the boundary condition for the frequency analysis. The mode shape of the flex spline was confined along the axial and radial directions. The eigenvalue of the axial modes was approximately five times greater than that of the radial modes. Because the thick ness of the axial direction is greater than that of the radial direction. The first mode of the axial direction was mea sured by 3650 Hz, and the error between the measured and estimated of frequency was 5" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003684_j.matpr.2021.05.466-Figure25-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003684_j.matpr.2021.05.466-Figure25-1.png", "caption": "Fig. 25. Circumferential stress for aluminium alloy.", "texts": [], "surrounding_texts": [ "Figs. 16, 17, 18, and 19 shows the result of longitudinal stressinduced in Elliptical, Hemispherical, Torispherical, and Plain formed head. Fig. 23. Circumferential stress for low carbon steel. Fig. 24. Longitudinal Stress for alluminium alloy. Fig. 26. Circumferential stress for gray cast iron. Fig. 27. Longitudinal stress for gray cast iron. Fig. 29. Longitudinal stress for stainless steel. Fig. 30. Longitudinal stress for titanium alloy. Fig. 31. Circumferential stress for titanium alloy. From graph Fig. 20, Stress-induced in Hemispherical head is minimum, while in Plain formed head, stress-induced is maximum. Therefore, for the taken conditions, 8 bar capacity pressure vessels and 24 L, the hemispherical end is better." ] }, { "image_filename": "designv11_22_0000261_s11548-019-02009-w-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000261_s11548-019-02009-w-Figure3-1.png", "caption": "Fig. 3 Sketch of the segmentwise robot construction. Starting with the fixed end-effector pose, the positions Pi of the robot segments are iteratively selected on spheres (a), which results in a search tree (b). Each selected point fixes two joints and the last remaining joint can be set arbitrarily afterward (c)", "texts": [ " Segmentwise robot construction (SRC) Arobot armconsists of a sequence of links connected by joints. Typically, the robot base is fixed and a certain pose of the end-effector is desired. Solving the inverse kinematics problem then results in a set of joint angles realizing the respective pose. In our approach, we do not fix the position or orientation of the base but iteratively select the pose of each link, starting from the predefined end-effector pose. Consider five points Pi , i = 1, . . . , 5, representing origins of coordinate frames according to the Denavit\u2013Hartenberg rules (Fig. 3). The position of P4 and P5 is defined by the ultrasound probe\u2019s pose. The other three points are subject to our optimization and include the position of the robot base (P1) and its orientation (vector from P1 to P2 and arbitrary rotation around this vector). In joint space, each of the points corresponds to the selection of two joint angles (Fig. 3c). If the points P5 to Pi (i < 5) are already fixed, the next free point Pi\u22121 is located on a sphere around Pi with a radius equal to the physical distance of these two points. We evaluate 36 candidate points on the sphere with respect to the objective value and select the best one. After obtaining an initial objective value for a tupel (P1, P2, P3, P4, P5), we search for better solutions in a depth-first manner. Branches of the search tree (Fig. 3b) are truncated, if they cannot lead to better solutions or would position the base outside of our search space. This approach is complete and optimal with respect to the discrete set of candidate points. Generally, a finer discretization improves the solution at the cost of the computational effort. A disadvantage of this approach is that we cannot explicitly control the LIFT angle of the robot, i.e., it is not possible to consider multiple angles in the optimization. Simulated annealing (SA) Simulated annealing is a generic stochastic approach which does not require calculation of derivatives and is therefore suitable for optimization of nonsmooth functions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003178_j.jsv.2015.06.037-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003178_j.jsv.2015.06.037-Figure15-1.png", "caption": "Fig. 15. The mode shapes of healthy rotating impeller\u2013shaft-bearing assembly: (a) the 2ND impeller dominated mode and (b) the 3ND impeller dominated mode.", "texts": [ " The linear model refers to an assembly with a 60 mm length of crack, but ignores the crack breathing effects. The nonlinear case involves the nonlinear crack breathing effects. As can be seen in Fig. 14(a), the 60 mm crack does not lead to significant changes in the response excited by EO1 traveling wave forces. For the forced response of EO2, the results exhibit a different characteristic from that of EO1. The response amplitudes given by the healthy assembly are approximately zero, due to the characteristics of 2ND impeller-dominated mode shapes, such as the one shown in Fig. 15(a). When such a mode is excited, the shaft bending deflection will be very small. When a crack exists, the response amplitudes are obviously enlarged due to the distortion of modes by crack. But their amplitudes are several orders of magnitude smaller than the response of EO1 excitation. Moreover, the crack breathing effects lead to the increase of the assembly's effective stiffness, and the response amplitudes decrease in some degree. Differing from the results of EO2, the nonlinear crack breathing effects remarkably change the response of EO3. The amplitudes of nonlinear case at some nodes are even larger than the linear ones, which may be due to the complex coupling between the 3ND impeller-dominated modes shown in Fig. 15(b) and other shaft-dominated modes. The transverse response of EO11 shown in Fig. 14(d) is consistent with the orbit shown in Fig. 13(b). The amplitudes of nodes between the two bearing nodes are quite small, whereas the responses of other nodes are much larger. The nonlinearity due to crack breathing effects also makes the response more complicated. Finally, the shift of response of the probe node versus crack length is presented in Fig. 16, where the results of linear and nonlinear models are included" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001291_j.proeng.2014.06.011-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001291_j.proeng.2014.06.011-Figure1-1.png", "caption": "Fig. 1 Experimental setup. Fig. 2 Two-dimensional double pendulum model. = ( ) 2( ) sin 1( ) + 2( ) 2( ) cos 1( ) + 2( )2 1( ) cos 1( ) + 2( ) 2 2( ) cos 1( ) + 2( )1( ) sin 1( ) + 2( )1( ) cos 1( ) + 2( ) 1( ) cos 1( ) 1 (3)", "texts": [ " The experimental task was to execute a kicking motion in which the ball is caught at the instep (the portion centred on the dorsum of the foot leading to the ankle). Each participant was asked to warm up, and then, with an ad libitum running start, to kick a soccer ball that had been set down towards a goal 10 m away, using the dominant leg at full force. Imaging was performed using 10 infrared cameras (Vicon Motion Systems, Oxford, UK); three-dimensional (3D) coordinate data for each body part (16 anthropometric points with reflective markers attached) during the kicking motion were collected at 250 Hz (Fig. 1). The stationary coordinate system was defined as a right-handed system in which the x-axis is the direction orthogonal to the horizontal kicking direction at the start of the task, the y-axis is the horizontal kick direction at the start of the task, and the z-axis is the vertical direction. The data, including the extrapolated points, were smoothed using a fourth-order phase-shift-free Butterworth digital filter to determine the optimum cut-off frequency (20 Hz) (Winter, 2004). A force platform (Kistler,Winterthur, Switzerland, Type 9287) was installed beside the ball, and the ground reaction force at the point of contact with the supporting leg was measured at a sampling frequency of 1000 Hz" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001554_0959651815595909-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001554_0959651815595909-Figure3-1.png", "caption": "Figure 3. The sideslip angle of tire in dynamic model.", "texts": [ " In this model, velocity vector is not along the tire motion direction, and the sideslip angle is formed from their difference. In fact, when the driver commands the vehicle in a road bend, velocity vector of vehicle center is not of the same direction with the wheel motion, and they make an angle with each other, which is called the sideslip angle. If the angle is large, the tires will be rubbed. In equation (6), ar and af are the sideslip angles of rear and front wheels of vehicle dynamic model, respectively, that are attained from the following equations, according to Figure 3 af = d uvf ar = uvr \u00f07\u00de According to Figure 3, d is the wheel steering angle with longitudinal axis, and uvf and uvr are the angles that velocity vectors of front and rear wheels have with at UNSW Library on August 21, 2015pii.sagepub.comDownloaded from longitudinal axis, respectively, and are obtained from the below equations tan uvf= vy +Lf _c vx tan uvr = vy Lr _c vx \u00f08\u00de where Lf and Lr are the distances between front and rear tires with vehicle gravity center, respectively, and vy is the vehicle velocity along the lateral axis. By tangent approximation of small angles and considering vy = _y, the following equations will be achieved uvf= _y+Lf _c vx uvr = _y Lr _c vx \u00f09\u00de Replacing the above values in equation (7), and by approximation of _y= vxb for small sideslip angles, the subsequent equation is obtained af = d b Lf _c vx ar = b+ Lr _c vx \u00f010\u00de where _c is a function of longitudinal velocity vx, and d is the wheel steering angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001615_lars-sbr.2015.33-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001615_lars-sbr.2015.33-Figure1-1.png", "caption": "Fig. 1. Global and vehicle reference frames. A denotes the global reference frame and B the vehicle\u2019s. The distance between both frames is represented by \u03be.", "texts": [ " A quadcopter has four propellers located at the ends of each segment of a cross-shaped frame. This is an underactuated system because it has four control inputs for its 6 degrees of freedom. There is one input that produces the total thrust 1www.ardrone2.parrot.com 2http://gazebosim.org/ 978-1-4673-7129-2/15 $31.00 \u00a9 2015 IEEE DOI 10.1109/LARS-SBR.2015.33 73 provided equally by each propeller in the direction of the vehicle\u2019s Z axis. Also, there are three angular inputs that determine the roll (\u03c6), pitch (\u03b8), and yaw (\u03c8). A representation of the global and vehicle frames is shown in figure 1. In this paper the mathematical model of a quadcopter is based on the equations described by Jirinec [11]. The global and vehicle frames are related by the rotation matrix: R = \u23a1 \u23a3 c\u03b8c\u03c8 c\u03b8s\u03c8 \u2212s\u03b8 s\u03c6s\u03b8c\u03c8 \u2212 c\u03c6s\u03c8 s\u03c6s\u03b8s\u03c8 + c\u03c6c\u03c8 s\u03c6c\u03b8 c\u03c6s\u03b8c\u03c8 + s\u03c6s\u03c8 c\u03c6s\u03b8s\u03c8 \u2212 s\u03c6c\u03c8 c\u03c6c\u03b8 \u23a4 \u23a6 (1) Where c\u00b7 and s\u00b7 are abbreviations for cos \u00b7 and sin \u00b7 respectively. The vehicle\u2019s position, linear, and angular velocity are represented by the vectors: \u03be = [ x y z ] T V = [ u v w ] T \u03c9 = [ p q r ] T The equations of motion are: \u03be\u0307 = R\u22121V (2) F = mV\u0307 + \u03c9 \u00d7mV (3) \u23a1 \u23a2\u23a3 \u03c6\u0307 \u03b8\u0307 \u03c8\u0307 \u23a4 \u23a5\u23a6 = E\u22121 \u23a1 \u23a3 p q r \u23a4 \u23a6 (4) R \u23a1 \u23a3 0 0 mg \u23a4 \u23a6\u2212 \u23a1 \u23a3 0 0 T \u23a4 \u23a6 = m \u23a1 \u23a3 u\u0307+ qw \u2212 rv v\u0307 + ru\u2212 pw w\u0307 + pv \u2212 qu \u23a4 \u23a6 (5) Where: \u2022 m: vehicle\u2019s mass \u2022 E: skew-symmetric matrix \u2022 T: thrust equally created by the four propellers The AR" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure4.3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure4.3-1.png", "caption": "Fig. 4.3 The shortest rotation (\u03b1) that brings a parallel to b is about an axis perpendicular to a and b", "texts": [ "1, but now assume that we have a targeting device that can be controlled with a quaternion. In other words, the zero quaternion describes the orientation where the targeting device is pointing straight ahead ([1, 0, 0]). Task: What quaternion would be needed to describe the target orientation, if the target is in an arbitrary location (x, y, z)? Solution: To answer that question, one can make use of the fact that the shortest rotation that brings a vector a into alignment with a vector b is a rotation about the direction perpendicular to a and b (see Fig. 4.3). n = a \u00d7 b |a \u00d7 b| (4.29) by an angle equal to the angle \u03b1 between the two vectors \u03b1 = arccos ( a \u00b7 b |a||b| ) . (4.30) Given the rotation axis and angle, the most convenient way to represent that rotation is the quaternion vector qadjust = n \u2217 sin(\u03b1/2) . (4.31) The corresponding algorithm is implemented in skin.vector.q_shorte st_rotation. For example, if the target moved along an \u221e loop on a screen in 10m distance, the orientation of the following targeting device could be calculated with the following code: Code: C4_targeting" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002204_0954406219900219-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002204_0954406219900219-Figure2-1.png", "caption": "Figure 2. Parameters of the finger seal laminates.", "texts": [ " The finger laminates are staggered and indexed in such a way that the fingers of one laminate cover the interstices between the fingers on the adjacent laminate, which blocks axial leakage, as shown in Figure 1(b) and (c). In addition, the flexible fingers can bend radially to accommodate the rotor excursions and the relative growth of the seal and rotor that result from rotational forces and thermal mismatch without damaging the integrity of the seal, which also gives the FS a longer life than conventional labyrinth seals. The FS structure used in this paper contains fourth finger laminates with arc-moulded beams. The main structure parameters and their values are shown in Figure 2 and Table 1, respectively, where Db is the diameter of the finger base, Df is the diameter of the upper finger foot, Di is the inner diameter of the FS, Dcc is the diameter of the finger beam arcs\u2019 centres, Rs is the arc radius of the finger beam, Is is the width of the gap between fingers, Ib is the width of the finger, Lst is the length of the finger, is the finger repeat angle, 0 is the finger foot repeat angle and b is the thickness of the laminate. Table 2 shows the material of the fingers (GH605) and rotor (K477) used in this paper, whose properties17 are similar to those of AMS5537 and Mar-M247, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure17-1.png", "caption": "Fig. 17. Stress analysis results of the pinion with two contact points (unit: MPa): (a) von Mises stress; (b) shear stress.", "texts": [ " For its stress analysis, the contact unit size, material properties, boundary conditions and load application are the same with gears with single contact point. The numbers of units and nodes are 378644 and 1608008, respectively. Finite element model of internal gears with two contact points is shown in Fig. 15. The analysis results in Fig. 16 show that the maximum contact stress of internal gears with two contact points is 616.88 MPa. The maximum von Mises stress and shear stress of the pinion with two contact points in Fig. 17 are 417.21 MPa and 133.82 MPa, respectively. The maximum von Mises stress and shear stress of the internal gear with two contact points in Fig. 18 are 347.72 MPa and 165.97 MPa, respectively. Obviously, two peak stress regions can be found on tooth profiles, which are also corresponding to two contact points. The stress distribution area is relatively concentrated and it has the trends of expanding towards the tooth root direction. The maximum contact stress of tooth profiles with two contact points is 56" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure11.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure11.12-1.png", "caption": "Figure 11.12 Direct-mounted cutter", "texts": [ " This type of collet chuck is located, held and driven in the machine spindle in the same way as the previously mentioned types. Large face mills are mounted directly on the spindle nose. To ensure correct location and concentricity, a centring arbor with the appropriate international taper is held in the spindle by the drawbar. The diameter on the end of the centring arbor locates the cutter, which is driven by the spindle keys through a key slot in the back face of the cutter. The cutter is held in position by four screws direct into the spindle nose, Fig.\u00a011.12. is securely held by the drawbolt. Spacing collars are slipped on the arbor, again ensuring that all faces are clean and free from dirt and metal cuttings. The cutter is positioned and spacing collars are added, together with the running bush, to make up the length of the arbor. The arbor nut is then screwed in position \u2013 hand-tight only. The arbor support is now positioned on the overarm so that it is central on the running bush and is then clamped in position. The arbor nut can now be tightened with the appropriate spanner" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001223_iciinfs.2014.7036598-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001223_iciinfs.2014.7036598-Figure1-1.png", "caption": "Figure 1: Second level adaptation.", "texts": [ ",\u00b5p] \u2208 Z are the number of points between maximum and minimum values of each parameter (including \u03b8p min and \u03b8p max), then the total number of models is given by N = \u00b51 \u00d7 \u00b52..... \u00d7 \u00b5p. The space of the models based on the arrangement given above will be the cartesian product of these sets, denoted as, S = [\u03b81 min, ...,\u03b81 max]\u00d7 [\u03b81 min, ...,\u03b81 max]\u00d7 ... ... \u00d7 [\u03b8p min, ...,\u03b8p max]. (17) The adaptation of parameters of the models as well as the movement of the virtual model towards the actual value of the plant can be seen in Figure 1. Combination of models here implies allocation of some adaptive weights to each of the model based on their identification error. The weighted summation of all the models provides the estimated plant parameter vector to be used in control input equation as per the certainty equivalence principle. It could be noted here that convergence of virtual model is faster than the actual models. In Figure 1, S is the bounded set for parameter vectors, \u03b8\u03021, \u03b8\u03022, \u03b8\u03023, \u03b8\u03024 are the parameters corresponding to four models and \u03b8v is the virtual model created by combination of all the models. \u03b8v(t) = \u2211 N i=1 wi(t)\u03b8i(t) (18) Following conditions must be fulfilled by wi(t), \u2022 The contribution from a model can never be neg- ative, i.e. wi \u2265 0. \u2022 Sum of contributions from all the models must be unity i.e., \u2211 N i=1 wi(t) = 1. Subtracting \u03b8 from both sides of (18) and using above properties, it can be shown that \u2211 N i=1 wi(t)\u03b8i(t) = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000474_marss.2019.8860956-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000474_marss.2019.8860956-Figure2-1.png", "caption": "Fig. 2. (a) Depiction of microrobot tumbling movement (b) Motion lapse of tumbling movement over aluminum surface. Scale bar is 1 mm.", "texts": [ " Due to orientation differences between the microrobot\u2019s magnetic alignment and that of the external actuating field, seen in Fig. 1(b), a magnetic torque is induced on the robot: ~Tm =Vm ~M\u00d7~B (1) Eq. 1 describes the general working principle of this torque, where Vm is the magnetic volume of the robot, ~M is the magnetization of the robot, and ~B is the external magnetic field strength. Under a time-varying rotating magnetic field, the torque causes the microrobot to undergo a forward tumbling motion. Fig. 2 depicts the resulting motion of the microrobot over a flat surface with no-slip conditions. While a cylindrical design would result in a more uniform rolling motion, standard MEMS fabrication techniques limit the thickness of the polymer block. A thin, rolling cylinder tends to tip over on rough surfaces, where small disturbances upset the stability of the cylinder, making a tumbling block better suited for biomedical applications. Fig. 3 summarizes the entire fabrication and magnetization procedure for the new microrobot design" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002085_1350650116681941-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002085_1350650116681941-Figure10-1.png", "caption": "Figure 10. Experimental instrument and friction samples: (a) Universal Micro-Tribotester-2 (UMT-2), (b) friction samples.", "texts": [ "91 10 7 4.43 F0.01\u00bc 29.46 4 Sbi 8.16 10 7 3 2.72 10 7 6.14 F0.05\u00bc 9.28 ( Sci 2.70 10 8 3 8.99 10 8 2.03 F0.1\u00bc 5.36 \u2013 Svi 6.18 10 7 3 2.06 10 7 4.65 F0.20\u00bc 2.9 4 e 1.33 10 7 3 4.43 10 8 Sum 2.41 10 6 15 ** represents highly significant, * denotes significant, ( represents having influence, 4 denotes having little influence. Figure 9. Effects of the selected parameters on friction coefficient f. at Purdue University Libraries on November 30, 2016pij.sagepub.comDownloaded from Micro-Tribotester (UMT-2) (Figure 10(a)). A pin moves reciprocally on a stationary flat surface in the experiment process. Surface topography of the pin is similar to the relevant flat, and equivalent surface parameters are owned for these two surfaces. Both the friction samples are completely immersed in lubricating oil (LG68 rail oil), and the system is maintained at room temperature and atmospheric pressure. Instantaneous friction force, contact load, and friction coefficient are measured by the triaxial forces sensor in UMT-2. All the pins and flats are made of quenched HT250 cast iron. The pin diameter is 5mm and the flat size is 30mm 40mm 6mm as shown in Figure 10(b). Table 8 lists the surface parameters of the friction samples, which are numbered as 1, 2, 3, and 4. These friction samples are selected according to the corresponding simulation results, and the friction force of the relevant simulation models decreases from 1 to 4. Before the tests, all the samples were screened using an optical microscope to ensure that the surface inclination does not influence the load. All samples were cleaned in acetone for 15min using an ultrasonic cleaner. An initial preload of 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000902_ilt-01-2015-0003-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000902_ilt-01-2015-0003-Figure3-1.png", "caption": "Figure 3 Schematic of contact of two surfaces", "texts": [ " On the running-in behavior of rough surface Jun Cheng, Xianghui Meng, Youbai Xie and Wenxiang Li Industrial Lubrication and Tribology Volume 67 \u00b7 Number 5 \u00b7 2015 \u00b7 468\u2013485 \u25cf For each small region, the asperities have spherical summits of same radius and there is no interaction with each other. In each small region, the asperity heights of an arbitrary cross-section taken along the x direction of piston ring surface and cylinder liner wall are z1(x) and z2(x). According to Pawlus and Zelasko, 2012, the heights of the equivalent surface roughness z(x) z1(x) z2(x) are shown in Figure 3. The standard deviation of asperity heights s, asperity radius and asperity density are adopted to evaluate the surface of piston ring and cylinder liner. To obtain the topography parameters of piston ring and cylinder liner, the spectral moments m0, m2 and m4 of the rough surface are calculated as follows (McCool, 1987): m0 AVG z2 (2) m2 AVG dz dx 2 (3) m4 AVG d2z dx2 2 (4) where AVG calculates the arithmetic average and m2 and m4 can be obtained by the finite difference method. Then the topography parameters , and s are decided as follows: m4 6 3 m2 (5) 0", " On the running-in behavior of rough surface Jun Cheng, Xianghui Meng, Youbai Xie and Wenxiang Li Industrial Lubrication and Tribology Volume 67 \u00b7 Number 5 \u00b7 2015 \u00b7 468\u2013485 A four-stroke spark-ignition engine is taken as an example to conduct the numerical simulation. The typical piston ring pack consists of top compression ring, scraper ring and one-piece twin-land oil control ring. The initial surface topography of piston rings is illustrated in Figure 10. Based on the real surface topography measured by the AFM, the profile of the contacting surface which is used in the model has been determined, as shown in Figure 3. The initial surface parameters of the piston rings and the cylinder liner for the sampling interval 10 m are listed in Table III. In this research, the height of the cylinder liner honing texture is close to the surface roughness, so the surface topography of cylinder liner is indicated by the surface roughness. The basic engine specifications and operating conditions are listed in Table IV. According to equation (23), the variation in dynamic viscosity of SAE 10W30 with temperature is obtained, as shown in Figure 11" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001268_s0263574714002653-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001268_s0263574714002653-Figure5-1.png", "caption": "Fig. 5. (Colour online) Geometric model of the second PKL.", "texts": [ " The different solutions of the inverse kinematic problem for the second PKL can be written as follows: \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u23a7\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23a9 \u03d522 = asin ( x\u2212x2 l3 ) or \u03d522 = \u03c0 \u2212 asin ( x\u2212x2 l3 ) \u23a7\u23aa\u23aa\u23a8 \u23aa\u23aa\u23a9 \u03d532 = acos ( T2\u221a (x1\u2212x)2+(z1\u2212z)2 ) + atan ( y2\u2212y z2\u2212z ) \u03d532 = \u2212acos ( T2\u221a (x1\u2212x)2+(z1\u2212z)2 ) + atan ( y2\u2212y z2\u2212z ) \u03d512 = atan ( y\u2212y2\u2212l3\u00b7s\u03d532c\u03d522 z\u2212z2\u2212l3\u00b7c\u03d532c\u03d522 ) (13) http://journals.cambridge.org Downloaded: 12 Mar 2015 IP address: 131.183.72.12 where \u03d512, \u03d532 are two rotation angles around x0, and \u03d522 is a rotation angle around x32 (see Fig. 5). The expression of T2 is given as follows: T2 = l2 2 \u2212 (y1 \u2212 y)2 \u2212 (z1 \u2212 z)2 \u2212 l2 3c\u03d52 22 2l3c\u03d522 . (14) Thus, it can be concluded that the inverse kinematic model for the two PKLs has four possible solutions. In the following, the static model of the manipulator is developed. It consists in determining the forces and moments transmitted by the joints of the two PKL. This model will be used to compute the orientation error of the RAF manipulator generated by the clearances in the passive joints" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003447_s0263574721000229-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003447_s0263574721000229-Figure13-1.png", "caption": "Fig. 13. Definitions of tool frame Ftool, joint 6 frame F6, and external payload frame Fe.", "texts": [ " Specifically, frame Fworld is defined by using three positions of the SMR (spherically mounted reflector) near the base of robot. Then, the vectors of axis-1 (joint 1) and axis-2 (joint 2) in the frame Fworld are acquired by using the CPA method. And frame F0 is defined as one that moves 400 mm (d1) from the origin of frame F1 in \u2013Z-direction. In this way, the geometric errors \u03b4\u03b11, \u03b4a1, \u03b4\u03b81, and \u03b4d1 are removed. The real transformation from frame Fworld to F0 is also given in Fig. 12. Ftool is established by using the three positions of the SMR (as shown in Fig. 13 where vectors v1 and v2 are acquired from the three positions of the SMR, Xtoolaxis is coincident with v1, and Ztool-axis is a cross product of v1 and v2) measured by the laser tracker with respect to frame F6. Frame F6is also defined according to ref. [6] as shown in Fig. 13 where the vectors of axis-5 and axis-6 in frame Fworld are acquired by using the CPA method. Therefore, the unknown parameters \u03b4\u03b16, \u03b4a6, \u03b4\u03b86, and \u03b4d6 are also eliminated from the identification models. The real transformation from frame F6 to Ftool is given in Fig. 13. The real transformation (changing with the thickness of counterweights installed on the end effector) from frame F6 to Fe is also given in Fig. 13. The experimental procedures are illustrated in Fig. 14. Step 1 has already been done in the experimental setup (Section 6.1). Then, step 2 is to generate 46,656 and 933,120 candidate configurations for kinematic and non-kinematic calibrations, respectively. These configurations are the same with those in Section 5.1. Step 3 is to select the optimized measurement configurations (by using the https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574721000229 Downloaded from https://www" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001755_978-3-319-26327-4-Figure14.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001755_978-3-319-26327-4-Figure14.5-1.png", "caption": "Fig. 14.5 Nonholonomic mobile robot", "texts": [ " In this section, the results from Sect. 14.3 are extended to includenonholonomic robots. That is, the hybrid automaton shown in Fig. 14.4 is reconsidered when the system is subjected to the nonholonomic constraint. For the nonholonomic systems, additional considerations are required to solve the regulation problem and the formation control problem due to the added complexities introduced by the nonholonomic constraint. In this section we consider a group of N networked nonholonomic systems. Consider the nonholonomic robot shown in Fig. 14.5. The equations of motion about the center of mass, C, for the ith robot are written as Fierro and Lewis (1998) q\u0307i = \u23a1 \u23a3 x\u0307ci y\u0307ci \u03b8\u0307i \u23a4 \u23a6 = \u23a1 \u23a3 cos \u03b8i \u2212di sin \u03b8i sin \u03b8i di cos \u03b8i 0 1 \u23a4 \u23a6 [ vi \u03c9i ] = Si(qi)v\u0304i, (14.15) where di is the distance from the rear axle to the robot\u2019s center of mass, qi = [xci yci \u03b8i]T denotes the Cartesian position of the center of mass and orientation of the ith robot, vi, and \u03c9i represent linear and angular velocities, respectively, and v\u0304i = [vi \u03c9i]T for the ith robot", " In fact, to obtain a posture stabilizing control policy it is necessary to use discontinuous and/or time-varying control laws (Luca et al. 2001). A technique which allows us to overcome the obstruction presented by Brockett\u2019s theorem is to apply a change of coordinates such that the input vector fields of the transformed equations are singular at the origin (Luca et al. 2001). This approach is carried out using a polar coordinate transformation, and the control law, once rewritten in terms of the original state variables, is discontinuous at the origin of the configuration space C. Consider again the robot shown in Fig. 14.5. Let \u03c1i be the distance of the point (xi, yi) of the robot to the goal point (xid, yid), let \u03b1i be the angle of the pointing vector to the goal with respect to the robot\u2019s main axis (lableled as XR in Fig. 14.5), and let \u03b2i be the angle of the same pointing vector with respect to the orientation error (Luca et al. 2001). That is \u03c1i = \u221a \u0394xi 2 + \u0394yi 2, \u03b1i = \u2212\u03b8i + arctan(2 (\u0394yi,\u0394xi)) + \u03c0, \u03b2i = \u03b1i + \u03b8i \u2212 \u03b8id, (14.18) where \u0394xi = xid \u2212 xi and \u0394yi = yid \u2212 yi. Then, the polar coordinate kinematics of mobile robot can be given as Luca et al. (2001) 348 H.M. Guzey et al. \u23a1 \u23a3 \u03c1\u0307i \u03b1\u0307i \u03b2\u0307i \u23a4 \u23a6 = \u23a1 \u23a2 \u23a3 \u2212 cos (\u03b1i) 0 sin(\u03b1i) \u03c1i \u22121 sin(\u03b1i) \u03c1i 0 \u23a4 \u23a5 \u23a6 [ vi wi ] . (14.19) From (14.19), it is observed that the input vector field associated with vi is singular for \u03c1i = 0 thus satisfying Brockett\u2019s theorem" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002338_tia.2020.2981572-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002338_tia.2020.2981572-Figure3-1.png", "caption": "Fig. 3. Structure finite element model.", "texts": [ " Therefore, an FFT analysis of the force density is performed in space and time, and the spatial order and the time harmonics of the force harmonics can be obtained, as shown in Fig. 2. It can be seen from Fig. 2 that the orders of the force density are mainly the 0th order, 6th order, 12th order \u2026 \u2026, which are related to the pole number. In order to analyze the structural vibration response, the exciting forces with a unit amplitude are applied to the stator of the motor. The three-dimensional (3-D) structure model of the motor can be seen in Fig. 3. The point A is the calculation point of the response. The materials parameters of the stator core include density p = 7850 kg/m3, Young\u2019s modulus E = 200 GPa, and Poisson\u2019s ratio v = 0.3. The effects of coil and varnish impregnation are not considered in the model. The structural dynamic simulation aims to obtain the vibration response for the motor. The harmonic analysis using the radial force frad,r with order r and unit amplitude is conducted. The simulation of the response excited by the radial force with orders Authorized licensed use limited to: University of Canberra" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001689_0040517514547210-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001689_0040517514547210-Figure3-1.png", "caption": "Figure 3. A geometric model of the right primary triangle.", "texts": [], "surrounding_texts": [ "A symmetric geometric model of spinning triangles in Siro-spinning is shown in Figure 1. It is easy to see that there are three spinning triangles, including two primary spinning triangles and one final spinning triangle, in Siro-spinning. Corresponding geometric models of the left and right primary spinning triangles and one final spinning triangle are shown in Figures 2\u20134, respectively. Here, w and h are the width and height of the primary triangle, is the apex angle of the primary triangle, /2 is the inclination angle of the yarn spinning tension, i is the angle between the ith fiber and the central fiber, and n is the fiber number in each fiber strand, that is, there are 2n fibers in the Siro-spinning triangle. O1 and O2 are the middle point of the nip line of each primary triangle, C1 and C2 are the twisting point of each fiber strand, and F1 and F2 are the spinning tension of each fiber strand. The final spinning triangle is shown in Figure 4. Here, C is the twisting point of the final spun yarn, F is the yarn spinning tension, which can be measured using a tension tester directly, and is the apex angle of the final triangle. According to force balance in Figure 4, we know F1 cos 2 \u00fe F2 cos 2 \u00bc F \u00f01\u00de F1 sin 2 \u00bc F2 sin 2 \u00f02\u00de Then, we have at BROWN UNIVERSITY on May 20, 2015trj.sagepub.comDownloaded from F1 \u00bc F2 \u00bc F sin 2 sin \u00f03\u00de In order to investigate the mechanical performance of a Siro-spinning triangle by using the FEM, the following assumptions are made.1\u20133 Assumption 1.1,3 The cross-section of all fibers is a circle with identical diameters. All fibers are gripped between the front roller nip and the twisting point. The velocity of fibers in the spinning triangle is constant, and the delivery velocities of fibers and yarn are the same. The stress\u2013strain behavior of the fibers follows Hooke\u2019s law for small strain. Assumption 2. 2 All the fibers are ideally packed in the yarn cross-section in concentric circular rings, where there are single core fibers at the center around which six fibers are arranged. Then, the number of fibers arranged in each layer is given by Nj \u00bc 1 j \u00bc 1 Nj \u00bc 6\u00f0 j 1\u00de j 2 ( \u00f04\u00de where Nj is the fiber number in the jth layer. Assumption 3. It is assumed that half of the apex angle is equal to the helical angle of the surface fibers of spun yarn.9 For an idealized yarn structure with a single-fiber core and closed packing of circular fibers, the surface helical angle can be calculated as follows:14 tan 2 \u00bc 2 r1T1 \u00f05\u00de Here, r1 is the radius of the final Sirospun yarn and T1 is the twist of the final Sirospun yarn. Assumption 4. The fibers in the primary spinning triangles are considered as three-dimensional (3D) elastic beam elements with tensile, compressive, torsional, and bending capabilities. The beam element, which exhibits a uniaxial line feature and has six structural degrees of freedom at each node, is chosen to represent the fiber. Meanwhile, all the degrees of freedom of the points of fibers on the front roller nip line are constrained.9 In the following, the spinning triangle of Ne40 Sirospun yarn will be analyzed. The material properties of cotton include the fiber elastic modulus E \u00bc 5Ntex, fiber shear modulus G \u00bc 1:351Ntex, fiber radius r \u00bc 0:008mm, fiber strand radius r2 \u00bc 0:0495mm, final Sirospun yarn radius r1 \u00bc 0:0715mm, and number of fibers in the Siro-spinning triangle 2n \u00bc 114. The other simulation parameters are listed in Table 1. In Table 1, M is the initial torque applied on two fiber strands produced by the fiber strand load, which can be calculated by M \u00bc 2F1r2 \u00fe 2F2r2 \u00f06\u00de at BROWN UNIVERSITY on May 20, 2015trj.sagepub.comDownloaded from" ] }, { "image_filename": "designv11_22_0003340_tia.2021.3058200-Figure25-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003340_tia.2021.3058200-Figure25-1.png", "caption": "Fig. 25 Demagnetization analysis at 0 elec. deg. rotor position with PM and 3 times of rated current excitation.\u00a0", "texts": [ " > REPLACE THIS LINE WITH YOUR PAPER IDENTIFICATION NUMBER (DOUBLE-CLICK HERE TO EDIT) < 9 through the other two teeth, which ensures that only little flux goes through the PMs. As shown in Fig. 22(d) and Fig. 23(a), when the rotor positions of the 18/17-pole IFC-BFPMM and 18/13-pole IFC-DSPMM are 270 elec. deg. and 0 elec. deg. respectively, their analyzed PMs have potentially the risk of demagnetization, because their directions of circumferential flux density are opposite to those of PMs. Further, the circumferential flux density distributions of the two machines generated by PM and 3 times of rated current excitation are shown in Fig. 24 and Fig. 25. The used magnet material is N35H. Because their working temperature may reach 120 \u2103, the demagnetization of the PMs will occur with the circumferential flux density lower than 0.35 T. In the 18/13-pole IFC-DSPMM, the area of the analyzed PM where the circumferential flux density is lower than 0.35 T is much smaller, which indicates that this machine has a better anti-demagnetization capability. An 18/13-pole IFC-DSPMM is manufactured and tested to verify the foregoing 2D FE analysis, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001703_epepemc.2014.6980569-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001703_epepemc.2014.6980569-Figure1-1.png", "caption": "Figure 1. Construction of an alternator of a car", "texts": [ " Nowadays special generators have already appeared on the market and built in cars, the so called starter-generators, as it is described in paper [1]. These are suitable to start the internal combustion engine as a starter motor and afterwards to operate as a generator. 1 housing, 2 stator, 3 rotor, 4 transistor voltage regulator with brush holders, 5 collector rings, 6 rectifier, 7 fan [2] In order to be able to simulate the behaviour of the voltage regulator we should know the construction and the operation of the entire system. The alternator basically is a synchronous generator as it is presented in Fig. 1. and described in many references e.g. [2]. It can be observed that the housing contains a rectifier and a voltage regulator, nearby other elements as brushes, fan, etc.. PEMC 2014 648 II. SIMULATION MODEL OF AN ALTERNATOR To create the simulation model of the alternator we started from a simplified version. An improvement of this model has been carried out by introducing into the model the followings: - the saturation of magnetic flux and - the effect of rotational speed of the shaft on the internal impedance of the synchronous generator" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001040_j.jappmathmech.2015.04.001-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001040_j.jappmathmech.2015.04.001-Figure2-1.png", "caption": "Fig. 2.", "texts": [ "10) would hold: For each current value of x along the trajectory of motion, Eq. (1.13) was solved for T(x) with the block-diagonal matrix I.M. Anan\u2019evskii, N.V. Anokhin / Journal of Applied Mathematics and Mechanics 78 (2015) 543\u2013550 549 a ( nd the controlling functions ui = (ai, i(T)xi) (i = 1, 2) were calculated from formula (1.15) and were then substituted into the complete non-linear) equations of motion of a three-link pendulum. 5 a A f R 1 2 3 4 5 6 7 8 9 50 I.M. Anan\u2019evskii, N.V. Anokhin / Journal of Applied Mathematics and Mechanics 78 (2015) 543\u2013550 Figure 2 shows the phase trajectories of motion of the first, second and third links. The thin lines are graphs of \u0307i( i), and the thick lines re graphs of \u0307i( i). Graphs of the controlling functions u1(t) and u2(t) are represented in Fig. 3 by thick and thin lines, respectively. It can be seen that the proposed control law transfers the three-link pendulum to an upper unstable equilibrium position. cknowledgements This research was supported by the Russian Foundation for Basic Research (14-01-00356a and 14-01-00476a) and by the Programme or the Support of Leading Scientific Schools (NSh-2710" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001928_1.g001577-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001928_1.g001577-Figure2-1.png", "caption": "Fig. 2 Control devices of QTW-UAV.", "texts": [ "Wing tilt is not considered a control input because it is used to establish the QTW\u2019s operating configuration, similarly to the landing flaps of conventional fixed-wing aircraft. The pilot sets the wing tilt angles at a priori designated tilt angles [6]. These selectable wing tilt angles and the flight configurations are determined as shown in Table 1. Note that McART3\u2019s wing tilt angle and the four flap (flaperon) deflections are all zero in the clean (or cruise) configuration, whereas the front and rear flaps are down to 25 and 15 deg, respectively at \u201c0 deg (flaps down).\u201d Figure 2 depicts the layout of the control devices of theQTW-UAV. In addition to the four thrusters (\u03b4pw1, \u03b4pw2, \u03b4pw3, and \u03b4pw4), the rudder \u03b4rud and the flaperons (\u03b4fl1, \u03b4fl2, \u03b4fl3, and \u03b4fl4) of each wing are available as control surfaces. These actual control inputs are converted into virtual inputs for control design: \u03b4flelv, \u03b4pwelv, \u03b4flail, \u03b4pwail, \u03b4flrud, and \u03b4pwrud. In high-tilt-angle flight configurations, the pitching moment is generated primarily by \u03b4pwelv. In the same way, the rolling moment is generated primarily by \u03b4pwail, and the yawing moment is generated by \u03b4flrud" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000597_0954406219890368-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000597_0954406219890368-Figure6-1.png", "caption": "Figure 6. Radial motion.", "texts": [ " As a result, the displacement at a certain direction is a combined effect of the deformations of external loading force, centrifugal forces, and gyroscopic moments. Since it is difficult to measure the centrifugal forces and gyroscopic moments, the displacement due to the external force should be separated. When the spindle rotates, the radial motion of the spindle is measured by two displacement sensors positioned 90 apart. The displacement sensors measure the X and Y displacement of the axis of rotation to generate a red curve, as shown in Figure 6. The three dotted circles are the largest inscribed circle, the leastsquare circle, and the smallest circumcircle from the inside to the outside. As the spindle speed increases, the centrifugal forces and gyroscopic moments increase. Considering the isotropy characteristics of the centrifugal forces and gyroscopic moments, the radius of the radial motion increases but the center remains unchanged. Namely, the displacement of the center of the radial motion is just the deformation due to the external loading force" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002328_j.mechmachtheory.2020.103873-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002328_j.mechmachtheory.2020.103873-Figure7-1.png", "caption": "Fig. 7. Transverse section of gear blank of face-hobbed and face-milled hypoid gear.", "texts": [ " Furthermore, the machine settings for those two types of hypoid gear were designed to obtain the same size and direction of contact ellipse. The transmission error of face-milled hypoid gear is close to that of face-hobbed gear. The mesh model of face-hobbed and face-milled hypoid gear were shown in Fig. 6 . Because of the same assembly parameters, the face-hobbed and face-milled hypoid gear share the same mesh model. D wp and D wg denote the axial mounting distance, respectively. E is the offset. represents the shaft angle. The transverse section of gear blank of face-hobbed and face-milled hypoid gear are shown in Fig. 7 . The tooth trace line and 2D blank of pinion and wheel are different. For face-hobbed hypoid gear, the generatrix of face cone, pitch cone and root cone are parallel with each other. The standard shrinkage is selected for face-milled gear, so the generatrix of those three cones are intersected at one point. According to the mesh model, the finite element analysis model for face-hobbed and face-milled hypoid gear were demonstrated in Fig. 8 . The Poisson\u2019s ratio and Young\u2019s modulus are 0.298 and 209 GPa" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure4.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure4.6-1.png", "caption": "Fig. 4.6 a Example of an object that could be best grasped with a variable pitch angle of the gripper. The grasp on this object would be more robust. b Estimation of a plane in the grasp region. The plane estimation works, because only the upper side of the object is scanned due to the single point of view of the sensor. c The approach vector of the gripper coordinate system has to be aligned with the normal of the plane to ensure best grasping", "texts": [], "surrounding_texts": [ "So far, objects may be grasped from a pile or a bin and may be transferred to a target location coarsely defined, e.g., using the joint angles of the robot. For industrial applications or service robotics, it is often essential to place the grasped objects at a predefined position in a desired orientation. In these cases, model data of the objects are required. Employing the gripper pose estimation approach presented in Sect. 4.1 has the drawback that the pose of the grasped object w.r.t. the gripper (denoted as grasp pose) is unknown. Regarded from a different point of view, it has the essential advantage that only the parameters needed for grasping are determined prior to grasping. Thus, 4.3 Bin-Picking Application\u2014Grasp Pose Estimation 47 the grasp motion could be initiated earlier; idle times of the robot could be reduced and therefore pick and place cycle times could be minimized. Nevertheless, to be able to place the grasped object at a desired pose, the grasp pose has to be determined before the object is placed. In this section, approaches are presented that estimate the grasp pose of the object. This happens either using the sensor data used for gripper pose estimation, or during the transfer motion after the object has been grasped, using force/torque and acceleration sensors. In fact, by estimating the coordinates of the center of mass of the object and its principal axes of inertia, a finite set2 of pose hypotheses can be obtained. The next section describes a technique for grasp pose estimation based on visual data. Then, a procedure to estimate the inertial parameters of an object which has been graspedwill be presented. The derivation of robust pose hypotheses from inertial parameters is addressed and strategies to deal with pose ambiguities are proposed." ] }, { "image_filename": "designv11_22_0003489_ieeeconf49454.2021.9382622-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003489_ieeeconf49454.2021.9382622-Figure1-1.png", "caption": "Fig. 1. Tower-type Cooking Robot", "texts": [ " The operation plan in the recipe analysis algorithm uses a motion library that defines the basic operation of the system and can be executed by appropriately modifying it. A series of experiments are conducted on a real machine to input human-created recipes and ensure that the operation is planned with the correct cooking procedures. In the next section, we will describe the performance and function of a real tower-type prototype of a cooking robot. In this section, we will describe the prototype of a tower type cooking robot. The full picture of this prototype is shown in Fig. 1 . The main body is made up of an aluminum frame with dimensions of 55 cm (width), 70 cm (depth), and 450 cm (height) including the upper material loading unit. This dimension fulfills the criteria of the acceptable dimensions. A tower-type cooking robot consists of independent units stacked high and functions as a single system. The ingredi ents are placed in special cooking containers and are trans ported accordingly. A polypropylene container is selected as the cooking container. The ingredients in the cooking containers are transported to each unit by the transport 456 Authorized licensed use limited to: East Carolina University" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003405_s10846-021-01342-0-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003405_s10846-021-01342-0-Figure4-1.png", "caption": "Fig. 4 Representation of collision margin rs", "texts": [ " The alternative arrangement of Fig. 2b has the same magnitude and altitude as the formation example of Fig. 2a, differing only by their angular positions, \u03b8i . As seen in Fig. 3, \u03b8i is the angular position of the i-th UAV in respect to the East axis at the horizontal plane. It distinguishes each UAV\u2019s target position within the perimeter of the circle. The constant rs defines the minimum distance between UAVs for which collision still does not happen. Its value is directly related to the UAV\u2019s dimensions, as depicted in Fig. 4. Indeed, it has to be the value of the largest length, l, of a UAV. The problem considered in this paper is to design a guidance system for a formation flight with collision avoidance amongst the UAVs. More specifically, the objective is to design a guidance system with the tracking characteristics of the PID-based solutions without their drawback of not addressing collisions. The proposed solution is intuitively described as to combine, in unique architecture, the \u201ccollision avoidance aware\u201d Dec-POMDP policy in emergency situations with the \u201caccurate\u201d PID policy in safe zones" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002179_icuas.2016.7502682-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002179_icuas.2016.7502682-Figure5-1.png", "caption": "Fig. 5. Hardware development for a quadrotor UAV", "texts": [ " Figures 4(b), 4(d), and 4(f) illustrate the angular velocity of the quadrotor during this maneuver. Noisy measurement data and the estimated value from EKF are presented along with the true path to show the performance and effectiveness of the EKF to reduce the error and improve noisy measurements. Position and velocity estimation errors for this example are presented in Figures 4(g) and 4(h). The quadrotor UAV developed at the flight dynamics and control laboratory at the George Washington University is shown at Figure 5(a). We developed an accurate CAD model as shown in Figure 1 to identify several parameters of the quadrotor, such as moment of inertia and center of mass. Furthermore, a precise rotor calibration is performed for each rotor, with a custom-made thrust stand as shown in Figure 5(b) to determine the relation between the command in the motor speed controller and the actual thrust. For various values of motor speed commands, the corresponding thrust is measured, and those data are fitted with a second order polynomial. Angular velocity and attitude are measured from inertial measurement unit (IMU). Position of the UAV is measured from motion capture system (Vicon). Ground computing system receives the Vicon data and send it to the UAV via XBee. The Gumstix is adopted as micro computing unit on the UAV" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002086_tjj-2016-0066-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002086_tjj-2016-0066-Figure16-1.png", "caption": "Figure 16: Comparison of pressure distributions with variable speed rotorseal.", "texts": [ " Accompany with the rotor motion, the corresponding high and low pressure regions in subsequent monitor point P7 are turned into the dotted rectangle regions in Figure 13. Extensive research in the whole cycle is shown in Figure 15, the correspondence advanced angle is the response of seal flow field to rotor vibration. For constant speed rotor-seal model, the advanced angle between the high pressure region and the minimum clearance region is objective and constant. For variable speed rotor-seal condition, the circumferential pressure distributions are plotted in Figure 16, with rotor in the same position P2. The highest pressure position in circumferential direction, represented by diamond, is not same in variable speed rotor-seal model. Pressure gradient provides flow motive power and pressure difference has the same effect of an aerodynamic force. The analysis of inner aerodynamic force produced by seal flow field is the second part of coupled interaction between rotor and seal. Differences in pressure distributions represent the variety of aerodynamic force, and the induced aerodynamic forces dependency on rotating speed are analyzed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001886_we.2008-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001886_we.2008-Figure2-1.png", "caption": "Figure 2. The figure shows the coordinates used for the formulation. Each gear is represented by a rigid body with its own set of coordinates.", "texts": [ "18 The design includes the physical dimensions of the gears, along with the gear mesh stiffness and the torsional stiffness associated with the intermediate shafts connecting each gearbox stage. Normally, there are two bearings in parallel supporting the planet gear, commonly known as upwind and downwind bearings. Because the bearings are represented by springs in the model in the current version of the code, the combined stiffness of the two bearings is used to represent the one bearing. That is, the equivalent stiffness of the bearings in parallel is calculated based on the individual stiffnesses of each bearing. Figure 2 shows the coordinates of the lumped-parameter model representation for a planetary stage and a parallel stage gearbox. The planet bearing loads and the simulation approach have been validated in the past with a transient test and a normal operation test. The validation results can be found in the study of Gallego-Calderon.19 Kb D 2 4 Kyy Kyz 0 Kyz Kzz 0 0 0 0 3 5 (2) The generator model used in the studies is a typical 3rd order DFIG model. The block diagram of the system is depicted in Figure 3, with a gray arrow over the vabc voltages that signifies the possible occurrence of a grid fault" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003515_s11249-021-01434-w-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003515_s11249-021-01434-w-Figure10-1.png", "caption": "Fig. 10 Slip speed distribution in the X direction (Load: 2\u00a0N, U\u2217 = u s \u2215u 1 )", "texts": [ " The values of c for the central stripe and inverse triangular slip patterns adopted in the calculation are 2440 and 2170\u00a0Pa, respectively, as extracted from Fig.\u00a05. The maximum pressures with the central stripe and inverse triangular slip patterns are higher than those in Tribology Letters (2021) 69:58 1 3 Page 7 of 8 58 the non-slip case, and their locations are close to the outlet of the slider. The maximum pressure along Y = 0 locates at the border of EGC coating region near the outlet for the central stripe pattern. This phenomenon corresponds to the variation of slip velocity at the border, as shown in Fig.\u00a010a, wherein the slip velocity reaches the maximum value and it drops rapidly to zero in the outlet, leading to a high pressure in that location. Figure\u00a010b illustrates the slip velocity distribution with the inverse triangle slip pattern. The slip velocity increases gradually from the inlet to the outlet and reaches the maximum at the rear boundary of the slider. This paper experimentally and numerically investigates the effect of oleophobic bearing surface on bearing load carrying capacity. The study was conducted with a twodimensional lubricated contact of a stationary slip slider surface and a moving non-slip plane. The results indicate that if slip occurs at the entire stationary surface, then the hydrodynamic load carrying capacity must be less than the Reynolds load support, regardless of the magnitude of the boundary slip" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003001_j.engfailanal.2020.105195-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003001_j.engfailanal.2020.105195-Figure2-1.png", "caption": "Fig. 2. Location of the crack in the bicycle frame.", "texts": [ " The use of miniature specimens also reduces the costs of fatigue tests [25]. The analyses were carried out for a bicycle frame used in gravity mountain biking. The bicycle frame is made of thin-walled tubes made of high-resistance steel alloyed with chromium and molybdenum. Fig. 1 shows the geometry of the analysed bicycle frame. After three years of use, fatigue cracks were detected at the connection between the top tube and the seat stays. The crack was initiated on both sides of the bicycle frame. Fig. 2 shows the failure location. The crack is located near the weld bead, and most of it is in the heat-affected zone. Fig. 3 shows the actual bicycle frame node with a visible fatigue crack. The crack is 10 mm long and is visible on the inner and outer sides of the tube. The crack is not directly on the edge of the welded joint, and it is offset from the edge by 1.5 mm. Visible corrosion can be observed on the inner side of the tube and not in the crack area. The tests aimed to identify the condition of the tested structures in accordance with the safety requirements for mountain bicycle frames" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001286_s2095-7564(15)30296-8-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001286_s2095-7564(15)30296-8-Figure8-1.png", "caption": "Fig. 8 Time history of acceleration and its time-frequency analysis ( V =130 km/h, L =40 mm)", "texts": [ " When the running speed of vehicle is 50 km/h and the size of wheel flat is 10 mm, the time history of axle box vertical accel eration and its time-frequency analysis are shown in Fig. 6. When the running speed of vehicle is 110 km/h, and the size of wheel flat is 10 mm, the time history of axle box vertical acceleration and its time-frequency analysis are shown in Fig. 7. When the running speed of vehicle is 130 km/h, and the size of wheel flat is 40 mm, the time history of axle box vertical acceleration and its time-frequency analy sis are shown in Fig. 8. When the running speed of vehicle is 150 km/h and the size of wheel flat is 70 mm , the time history of axle box vertical acceleration and its time-frequency analysis are shown in Fig. 9. From Figs. 6 and 7 , the W -R shock caused by wheel flat does not lead to the loss of W-R contact be cause the time history of axle box acceleration has no constant part. There is a single energy peak in the time-frequency image. But in Figs. 8 and 9, the time history of acceleration has a constant part, loss of W-R contact occurs, and multiple energy peaks ap pear" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002384_s12206-020-0325-y-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002384_s12206-020-0325-y-Figure10-1.png", "caption": "Fig. 10. Von Mises stress results for the initial and optimum models for steel and composite sides.", "texts": [], "surrounding_texts": [ "Two kinds of helical gears, an aramid/phenol and 100 % steel gears, were fabricated for vibration tests to investigate the damping effect of the composite part inserted between inner and outer steel parts. Force is applied to a tooth of two helical gears in a tangential direction by using a shaker, as shown in Fig. 11. The experimental setup for vibration test is depicted in Fig. 11. Force is imparted to the point A on a tooth surface and then acceleration is measured at the point B, as shown in Fig. 11, in order to check vibration transfer to the central shaft. Accelerations of aramid/phenol and 100 % steel gears are compared in Fig. 12. It is noticed from Fig. 12 that acceleration in a aramid/phenol hybrid gear decays a little more faster than 100 % steel gear. Fig. 13 shows FFT results of Fig. 12 in a frequency domain. From Figs. 12 and 13, it is verified that aramid/phenol composite region can absorb vibration. The variations in damping capability may be assessed by checking the logarithmic decrement ratio which is defined as the natural log of the amplitudes ratio between any two successive peaks and the damping ratio is then found from the logarithmic decrement relationship shown as follows [17]. 11 ln 1 d \u00e6 \u00f6 = \u00e7 \u00f7 - \u00e8 \u00f8n X n X (8) 2 2(2 ) dz p d = + (9) where \u03b4 denotes the logarithmic decrement ratio, X1 is the first cycle\u2019s acceleration output, Xn is nth cycle\u2019s acceleration output, n is a cycle number, and \u03b6 denotes the damping ratio. Here, n = 6 is applied. Table 5 also show that the damping ratio of the hybrid gear is higher than that of 100 % steel gear, based on calculations from Eqs. (8) and (9). Table 3. Optimum values of the design variables. Initial (mm) Optimum (mm) h 3 3 upw 3 2.25 downw 3 2.6 d 1.5 1.13 l 1 0.37 r (radius) 0 2 Table 4. Comparison of max von Mises stresses for initial and optimum models. Initial Optimum Composite part 118.0 94.0 20.24 %\u25bc Max von Mises stress (MPa) Steel part 813.0 697.0 14.25 %\u25bc * steels = 800 MPa, phenols = 100 MPa Table 5. Comparison of damping ratios for steel and hybrid gears. Steel gear Hybrid gear Improvement \u03b6 0.0387 0.0551 42.3 %\u25b2 Fig. 11. Experimental setup for vibration tests." ] }, { "image_filename": "designv11_22_0003398_s00466-020-01960-9-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003398_s00466-020-01960-9-Figure7-1.png", "caption": "Fig. 7 Residual stress fields for the evolving gas/metal interface for an additive buildup. Results shown are from a timestep 1, b timestep 50, c timestep 250, and d timestep 1000", "texts": [ " In the flat plate example, the high-fidelity multiphysics workflow shows a divot forming and moving as the laser scans; a distinct feature not captured by the part-scale process model. Through the inclusion of additional physics associated with laser interactions, and gas/fluid effects, the resulting cross sectional profile of the high-fidelity coupled workflow is able to much more closely match the experimental observation compared with the part-scale workflow. We performed similar comparisons for the 4-pass LENS additive build simulation. Figure 7 demonstrates the high fidelity thermal/fluid results along with the associated residual stress predictions. Figure 8 demonstrates the residual stresses predicted by the part-scale model for a similar build, again demonstrating that the multiphysics workflow yields a significantly higher residual stress prediction when compared to the part-scale model. Line plots comparing von Mises stress results for the part-scale model and the highfidelity multiphysics workflow at the build/plate interface along the center of the build are shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001137_ipec.2014.6869708-FigureI-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001137_ipec.2014.6869708-FigureI-1.png", "caption": "Fig. I. Structure of 8/1 0 hybrid stator pole type BLSRM.", "texts": [], "surrounding_texts": [ "reluctance motors (BLSRMs), 8/10 hybrid type and 12/14 hybrid type, are presented. The two BLSRMs have separate torque and suspending force poles. Due to the independent characteristics between the torque and suspending force, the torque control can be decoupled from the suspending force control. In this paper, suspending force control characteristics of the two hybrid BLSRMs are compared. Although the two hybrid BLSRMs have excellent decoupling characteristics, the suspending force control performance is dependent on the torque characteristics. In the suspending force control comparison, a simple PID and hysteresis current control methods are employed. To verify the validity of the control strategy, tests are executed to the two types, and the control characteristics of the two types are presented and compared.\nKeywords- Bearingless motors, force control, magnetic suspension, switched reluctance motorrs.\nI. INTRODUCTION\nMany modem industrial applications, such as high speed machine tools, turbomolecular pumps, centrifugal pumps, compressors, flywheel energy storage and aerospace need high or ultra-high speed machines. However, many problems may arise, when traditional mechanical bearings are used to bear the shaft of high or ultra-high speed machine. For instance, the mechanical bearing can cause increased frictional drag, thermal problems and heavy wear in high-speed motors, which may affect the efficiency, bearing life and maintenance of the machine. Further, the lubrication oil that is required by the mechanical bearings cannot be used in high vacuum, ultra high temperature and low temperature environments [1]-[2]. In order to solve these problems caused by the mechanical bearings, magnetic bearing motors are researched [3]-[4]. Magnetic bearing motors have the advantages such as friction-free, abrasion-free, seal-free, lubrication-free, high speed, high precision, long life, easy-to-implement active control, and so on. However, high power density is not easy to be implemented due to the complex structure. The separated controllers are also needed because the torque and suspension part are separated in the magnetic bearing motor, which will lead to high cost.\nBearingless switched reluctance motor (BLSRM) is developed on the basis of magnetic bearing motor and\n978-1-4799-2705-0/14/$31.00 \u00a92014 IEEE 994\nswitched reluctance motor (SRM). Hence, BLSRM not only extends theory and application of the bearingless motor, but also inherits high-speed performance and adaptability to harsh environment of SRM. Therefore, BLSRM can achieve the operation of high speed or ultra high speed [5]-[6].\nRecently, several structures of BLSRMs, such as 12/8 double winding type, 12/8 single winding type, 8/6 single winding type and Morrison type, are proposed [7]-[10]. But all the structures proposed in [7]-[10] are based on general SRM structure, so the available suspending force region is limited, and the torque control is coupled with the suspending force control. To realize stably suspension in these structures, complex mathematics equations have to be derived [11 ]-[ 14], and the mathematics equations are built up without considering the magnetic saturation, which increases the difficulty in controlling.\nIn order to expand the available suspending force region and reduce the control difficulty of conventional BLSRMs, an 8/1 0 hybrid stator poles BLSRM is proposed [6]. There are two types of stator poles in this motor: torque and suspending force poles. Compared with conventional BLSRM, the suspending force performance is improved and the air-gap is easier to control, but in this structure only half of the stator poles are used for the torque, output power density is very low. Moreover, the effect of torque current on the suspending force is somewhat large when the torque and suspending force windings are excited simultaneously.\nIn order to further improve the performances of BLSRM, a novel 12/14 hybrid stator pole type BLSRM is proposed [15]. The proposed structure also has separated torque and suspending force poles. Further, in the 12/14 type, short flux paths are taken and no flux reversal exists in the stator core. Compared to the 8/10 hybrid stator pole type BLSRM, the output torque is significantly improved and the air-gap is easier to control.\nIn this paper, the suspending force characteristics of the two BLSRMs are compared. In the suspending force control comparison, the BLSRMs mathematical model is not used in the control scheme, only a simple PI controller is used to regulate the speed of the proposed BLSRM, and two PID controllers are used to generate the desired suspending force commands to keep the rotor in the centre position. Based on the control scheme, the", "experimental system is constructed and experiments are performed to the two BLSRMs. Finally, the control characteristics of the two BLSRM are presented.\nII. THE NOVEL BLSRMS WITH HYBRID STATOR POLE\nA. 8110 Hybrid Stator Pole Type BLSRM\nFig. 1 shows the structure of 8/10 hybrid stator pole type BLSRM. Different from the conventional structures, two types of stator poles are included in this structure. PAl and PA2 are the torque poles of the phase A, and PRI and Pm are the torque poles of the phase B. The x direction suspending force is generated by currents which flow in the windings of the suspending force poles P,p and Pxn. That is, the current i,p in the Pxp stator pole generates positive x-direction suspending force, and the current ixn in the Pxn stator pole generates the negative x direction suspending force. Similarly, the suspending forces for the y-direction are generated by the currents iyp and iyn, which flow in suspending force poles Pyp and Pym respectively. Meanwhile, to get a continuous suspending force, the suspending force pole arc needs to be wider than one rotor pole pitch. In that way, the aligned area between the suspending force and the rotor pole is always the same which may decrease the effect of suspending force current to the torque.\nB. 12114 Hybrid Stator Pole Type BLSRM\nA novel 12/14 hybrid stator pole type BLSRM is proposed in Fig. 2. The proposed structure is similar to the structure of the 8/1 0 type. There are also torque and suspending force poles in the 12114 type. But in the 12114 type short flux paths are taken and no flux reversal exists in the stator core. Windings on the torque poles PA/, PA2, P A3 and PM are connected in series to construct phase A, and windings on the torque poles PRJ, Pm, PRJ and PR\ufffd are connected in series to construct phase B. Similar to the 8110 type, the windings on the suspending force poles Pxp' Pxm Pyp and Pyn are independently controlled to construct four suspending forces in the x- and y directions. In order to get the continuous suspending force, the suspending force pole arc is also selected not to be less than one rotor pole pitch.\nC. Torque and Suspending Force Characteristics a/the twoBLSRMs\nAs is well-known, the torque and suspending force windings are always excited simultaneously in BLSRMs. Hence, to obtain the characteristics of the proposed structure, finite element method (FEM) is employed to get the characteristics of the proposed structures.\nFig. 3 shows the torque profiles with fixed torque current (iA=2A) and various suspending force currents. In Fig. 3, PU is per unit. 1 PU stands for one rotor pole pitch. That is, 1 PU stands for 36 and 26 mechanical degrees in the 8110 and 12114 types, respectively. From Fig. 3, it can be seen that the effect of suspending force current on the torque is very small in the two structures. Therefore, the suspending force current has almost no effect on torque.\nFig. 4 shows the suspending force profiles with fixed suspending force current (iyp=2A) and various torque currents. As seen in Fig. 4, when the torque winding current is zero, the suspending force has excellent linearity with respect to the rotor position in the two motors. Moreover, the available suspending force region in the proposed two structures is the whole rotor pole pitch, which is wider than the suspending force in conventional BLSRMs. It also can be seen from Fig. 4 that the effect of torque current on the suspending force is somewhat large in the 8/10 type, while in the 12/14 type, it is small enough to be ignored when compared to the suspending force generated by the suspending force winding.\nFrom the above analysis, it can be seen that the proposed two structures cannot perfectly decouple the torque and suspending force control, but compared to conventional BLSRMs, the coupling effect is extremely reduced, especially in the 12/14 type, the torque control is almost decoupled from the suspending force control.\nIII. CONTROL SCHEME OF THE PROPOSED BLSRMS\nAccording to the analysis in Section II, a control scheme for the two BLSRMs is proposed in Fig. 5. As shown in Fig. 5, mathematics model is not used in the control scheme, only a PI type speed controller is adopted", "to regulate the motor speed, and two independent c1ose loop PID air-gap displacement controllers, one for x direction and the other for y-direction, are used to generate the desired suspending force commands Fx * and Fy * to keep the rotor at the center position. Further, in the control scheme, the actual current values of the exciting phase of the suspending force winding can be controlled through the hysteresis method according to the command current signals.\nFig. 6 shows the suspending force generation principle of the proposed BLSRMs. As shown in Fig. 6, if the winding on the suspending force pole Pxp is excited, the suspending force F:, in the positive x-direction can be generated. If the winding on the suspending force pole Pyp is excited, the suspending force Fy in the positive y direction can be generated. In the same way, if the windings on the suspending force pole P,p and Pyp are excited simultaneously, the force F will be generated, which is the synthesis force of Fx and Fy - Furthermore, the value and direction of the force F can be regulated by changing the values of the currents in the two windings. Therefore, when controlling values of the currents in the four suspending force windings, the desired resultant\nforce in any arbitrary direction and magnitude can be obtained to compensate for the unbalanced pull force caused by the non-uniform air-gap.\nAccording to the suspending force generation principle in the proposed BLSRMs, the force F, in any direction and magnitude, can be generated by the windings on the two poles, i.e. one each in the x- and y-directions. For instance, assume that F is the desired force, as shown in Fig. 6. In this case, because F is in the first quadrant, pole" ] }, { "image_filename": "designv11_22_0003413_j.jmbbm.2021.104412-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003413_j.jmbbm.2021.104412-Figure10-1.png", "caption": "Fig. 10. Mechanism of soft robot using air actuators: (a) The uniform bending deformation and (b) the different bending deformation.", "texts": [ " The tendency of the wave pattern with a 90\u25e6 pattern angle exhibits similar results with the lane-type pattern having a 45\u25e6 pattern angle; this may be due to the fact that the wave pattern consists of continuous 45\u25e6 lane-type patterns. Overall, the changes in pattern geometry prove to serve as another parameter in changing the mechanical properties of PDMS specimens, where both the tensile strength and hyperelasticity change due to the changes in pattern geometry. This section applies the micro-patterns to medical soft robots, enabling enhanced robot movement and delicate motion control, as illustrated in Fig. 10. One of the common actuators of soft robots in medical applications is the pneumatic actuator by vacuum or compressed air, and constant pressure is developed inevitably inside the soft materials (Tse et al., 2018). For example, Fig. 10a shows an example with constant pressure and a soft robot of uniform thickness, in which uniform deformations occur. However, existing soft robots have limitations in having to change the input pressure of the pneumatic actuator when the bending angles of the joints change. This indicates that each joint of the soft robot must have individual air actuators to control the motion of the robot, thereby compromising its inefficiency. For enhanced motion, different pressure should be applied. To enhance the motion of a soft robot subjected to constant pressure, the material properties of the soft robot can be varied, as shown in Fig. 10b by attaching the specimens with the surface patterns. This indicates that joint parts do not require individual air actuators, thus increasing the efficiency in controlling soft robot motions. Moreover, more delicate motion control can be achieved by engraving the surface patterns of appropriate pattern angles and pattern geometries at the joints of soft robots. To validate the concept of the application of the lane-type PDMS specimens to soft robots, experiments were conducted as illustrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003114_tmag.2015.2438872-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003114_tmag.2015.2438872-Figure2-1.png", "caption": "Fig. 2. Proposed SPMSM for the washing machine.", "texts": [ "2438872 conducted to obtain the deteriorated B\u2013H curve and iron loss data. The test was carried out over several levels of excitation field at several frequencies, and the results were compared with those of the conventionally stacked lamination method. In addition, the iron loss of the motor was analyzed through the finite-element method (FEM) using the examined material data. It was finally compared with the experimental result of the actual iron loss of the motor. An outer-rotor-type SPMSM, as shown in Fig. 2, consists of 48 poles-36 slots, and a ferrite magnet was used in the rotor. The diameters of the rotor and the stator are 283 and 265 mm, respectively. The stator is slinky laminated with a 1 mm cold-rolled steel sheet, and the rotor is made by stamping a galvanized cold-rolled steel sheet. The stack length of 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002058_1.4035203-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002058_1.4035203-Figure9-1.png", "caption": "Fig. 9 The finite element model: (a) structure\u2013fluid coupled model and (b) the structure model", "texts": [ " In fact, the self-excited vibration needs the servomotor to supply energy. If the shaft works at a relatively low speed, the stronger friction from the water-lubricated rubber bearing will make the driving torque of the servomotor fluctuate to a large extent. In this circumstance, the shaft speed is unable to keep stable and the selfexcited vibration will disappear. To find out the associated unstable modes and derive modal parameters for further analysis, a finite element model of the flexibly supported shafting system submerged in water, as shown in Fig. 9, was built with the software ABAQUS. Parameters of the model are listed in Table 1. The water and the shaft are modeled with 506,496 solid elements, and the supports are modeled with 9472 shell elements. The supports and the shaft are connected by springs. To consider the influence of water, the nodes at the surfaces between the shaft and supports are coupled with the associated nodes of water, which have a nonreflecting boundary. The outside-ends of support beams as shown in Fig. 3 are fixed, and the pretightened force Fp is 1700 N" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000459_978-3-030-29041-2_16-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000459_978-3-030-29041-2_16-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_22_0001649_j.engappai.2015.11.009-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001649_j.engappai.2015.11.009-Figure2-1.png", "caption": "Fig. 2. Inverted pendulum system.", "texts": [ " lim t-1 Z t 0 W \u03c4\u00f0 \u00ded\u03c4o1 \u00f043\u00de Also, W t\u00f0 \u00de is a positive function, this means that _W t\u00f0 \u00de is bounded for all time, so by Barbalat's lemma (Slotine and Li, 1991; \u00c5str\u00f6m and Bj\u00f6rn, 2013; Khalil and Grizzle, 1996; Shahnazi and Akbarzadeh-T, 2008), it can be demonstrated that limW t\u00f0 \u00de \u00bc 0 t-1 according to (43). That is, E t\u00f0 \u00de-0 as t-1. The proposed control scheme is depicted in Fig. 1. To evaluate the adaptive LS-SVR control scheme proposed in this paper, first, a simulation example for an inverted pendulum system under different types and level of disturbances is performed, and then is compared with the reference adaptive fuzzy controller proposed by Wang (1996). Example. Consider the inverted pendulum system in Fig. 2. Let x1 \u00bc \u03b8 rad\u00f0 \u00de and x2 \u00bc \u03b8=s\u00de, the second-order dynamic equation of the inverted pendulum is given by Khalil and Grizzle (1996): x1 : x2 : \" # \u00bc 0 1 0 0 x1 x2 \" # \u00fe 0 1 f x1; x2\u00f0 \u00de\u00feg x1; x2\u00f0 \u00deu\u00fed y\u00bc 1 0 x1 x2 \" # \u00f044\u00de where f x1; x2\u00f0 \u00de \u00bc 9:8 sin x1 mlx22 cos x1 sin x1 mc \u00fem l 4 3 mcos 2x1 mc \u00fem \u00f045\u00de where mc \u00bc 1kg and m\u00bc 0:1kg are the masses of the cart and the pole, respectively, andl\u00bc 0:5m is half length of the pole, and u is the applied force (control signal). The control object is to control the state x1 of the system to track the desired trajectory xd \u00bc \u00f0\u03c0=30\u00de sin \u00f0t\u00de if only the system output is measurable" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000654_tnano.2019.2960126-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000654_tnano.2019.2960126-Figure5-1.png", "caption": "Fig. 5. (a) Illustration of a Janus particle used as a microrobot; (b) Microscope image of the microrobots; (c) Schematic of magnetic control.", "texts": [ " Finally, the sample was developed and the microchannels were created (Fig. 4(b)). To ensure that the water solution with microrobots fully fill the microchannel, the microchannel\u2019s surface was treated using plasma cleaning (200 w and 3 min) to increased hydrophilicity. The microrobots are 10 \u03bcm Janus particles composed of polystyrene (PS) particles coated on one side with Chromium (Cr) and nickel (Ni) using electron-beam evaporation with the thickness of 20 nm and 100 nm respectively, as shown in Fig. 5(a) and (b). The Cr layer increases the absorption of Ni while the Ni layer allows the microrobots to be controlled via a rotating magnetic field. When a field is applied, the dipole m of the microrobots always aligns with the direction of the magnetic field; thus, as the field rotates, the microrobot will also rotate in order to maintain dipole alignment [12]. The motion of the rolling microrobot is manipulated by the external rotating magnetic field, which is supplied by a 3D Helmholtz coil system. The XY motion of the microrobots is controlled by adjusting the strength B, angle frequency \ud835\udf14, and orientation of the rotating magnetic field. The magnetic field used for 2D control can be expressed as: \ud835\udc35 = [ \ud835\udc35\ud835\udc5f sin(\ud835\udf03) cos(\ud835\udf14\ud835\udc61) \ud835\udc35\ud835\udc5f cos(\ud835\udf03) cos(\ud835\udf14\ud835\udc61) \ud835\udc35\ud835\udc5f sin(\ud835\udf14\ud835\udc61) ], (11) where \ud835\udc35\ud835\udc5f , \ud835\udf03, \ud835\udf14, and \ud835\udc61 represent the amplitude of the rotating magnetic field, the direction of motion, the rotating frequency of the field, and run time, respectively. The schematic for Eq. (11) is shown in Fig. 5(c). To validate the computation model, a series of experiments were conducted using a control system that combines the proposed algorithm with feedback control [13]. The control system is composed of a Helmholtz coil system, three power supplies, a National Instruments data acquisition (NI DAQ) device, a camera, a microscope, and a host computer (Fig. 6(a) and (b)). The Helmholtz coil is powered by three power supplies controlled via the DAQ device. Each coil pair generates a magnetic field of approximately 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001235_j.acme.2014.11.003-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001235_j.acme.2014.11.003-Figure7-1.png", "caption": "Fig. 7 \u2013 Dynamic model of an external gear pump.", "texts": [ " For the simulation model the trapezoidal changes of the gearing stiffness as in Fig. 6 has been considered. The described excitations cause rotational and translational vibrations of the gears in the gear pump and these are coupled through the meshing. To investigate the influence of these Please cite this article in press as: W. Fiebig, M. Korzyb, Vibration a Mechanical Engineering (2015), http://dx.doi.org/10.1016/j.acme.2014.11.0 excitation forces on the vibration of the gearwheels and dynamic loads in the external gear pumps the dynamic model shown in Fig. 7 has been considered. The following simplifications have been assumed: the vibrations of the gears are considered at a constant angular velocity of the drive shaft, inertial forces of the oil inside the pump can be neglected, there is no pressure overdue in trapped volumes and the axial forces are compensated, there are no forces from unbalance on the gears, tiffness along the line of action. nd dynamic loads in external gear pumps, Archives of Civil and 03 the manufacturing errors are negligibly small, the damping of torsional vibrations of the gears has a viscous character" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure16.8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure16.8-1.png", "caption": "Figure 16.8 Indirect extrusion", "texts": [ " The container is moved forward against a stationary die and the ram pushes the metal through the die. After extrusion, the container is moved back and a shear descends to cut off the butt end of the billet at the die face. The process is repeated with a new billet. As the outside of the billet moves along the container liner during extrusion, high frictional forces have to be overcome, requiring the use of high ram forces. Billet Dummy block Ram Container liner Container body Extrusion Die Pressure plate Figure 16.7 Direct extrusion Figure 16.9 Examples of extruded shapes In this process, Fig.\u00a016.8, the heated billet is loaded in the container which is closed at one end by a sealing disc. The container is moved forward against a stationary die located at the end of a hollow stem. Because the container and billet move together and there is no relative movement between them, friction is eliminated. As a result, D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 16 Primary forming processes 16 258 blacksmith. Hand tools are used to manipulate the hot metal to give changes in section and changes in shape by bending, twisting, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000690_morse48060.2019.8998714-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000690_morse48060.2019.8998714-Figure1-1.png", "caption": "Fig. 1. A schematic of omni robot.", "texts": [ " Software architecture for the robot control is constructed based on ROS to accompany the application of the SLAM algorithm for robot navigation task is given in section III. Hardware architecture and research results using GAZEBO and RVIZ are presented in section IV. Finally, a conclusion is mentioned in section V. II. ARCHITECTURE OF OMNI SYSTEM The Omni robot has four wheels which are 90 degrees apart. Oxy represents the global coordinate axis, the distance between wheels and the robot center was defined by d . The robot movement would be identified for the navigation stack. By the global coordinate as chosen in Fig. 1, the velocity of the robot contains three components: a linear velocity along x-axis and y-axis and angular velocity. Now, the robot coordinate vector is defined as [ ]T q x y \u03b8= and the velocity of the robot in the global coordinate could be obtained by taking the derivative of q . Moreover, the relationship between the velocity in the robot axis and the velocity in the world axis is described by the kinematic model of the robot: cos sin 0 sin cos 0 . 0 0 1 q v \u03b8 \u03b8 \u03b8 \u03b8 \u2212 = (1) From (1) we obtain the equations which would be used to program the robot navigation in ROS: cos sin sin cos x y x y x v v y v v w \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 = \u2212 = + = (2) where , ,x yv v and w are the velocity in the robot axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003349_s00170-021-06813-0-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003349_s00170-021-06813-0-Figure14-1.png", "caption": "Fig. 14 CAD model of in situ SLS cold plate including two copper-plated (0.8 mm) solid BN ceramic tubes (100mm\u00d7 185mm \u00d7 18 mm)", "texts": [], "surrounding_texts": [ "Active cooling prototype is represented by following equations that govern incompressible fluid flow of LM and heat transfer: The continuity equation (conservation of mass) \u2207:V \u00bc 0 \u00f04\u00de The momentum equation (Navier\u2013Stokes equations) \u03c1 \u2202V \u2202t \u00fe V :\u2207V \u00bc \u2212\u2207p\u00fe \u2207:\u03c4\u2212\u03c1g \u00f05\u00de Conservation of energy (the temperature distribution): \u03c1Cp \u2202T \u2202t \u00fe V :\u2207T \u00bc \u2207: k\u2207T\u00f0 \u00de \u00fe Q \u00f06\u00de where V is velocity of LM, \u03c1 is density of LM, T is temperature, p is pressure, t is cycle time, \u03c4 is deviatoric stress, g is gravity vector, \u03b2 is volumetric expansion coefficient, Cp specific heat at constant pressure, k thermal conductivity, Q is volumetric heat source. Ansys Fluent is used for computational fluid dynamics (CFD) simulation to provide approximate solutions to nonlinear partial differential governing equations. The equations are therefore converted into a set of algebraic equations solved through numerical techniques. Preprocessing efforts include adaptive tetrahedron mesh (Fig. 10), assigning material properties, and full contact is assumed at interface of plated tube and SLS aluminum alloy metallic block (100\u00d7185\u00d718mm3). Due to the limitation of the CFD software available by today\u2019s simulation methods, it only allows \u201cfull contact\u201d at interfaces, and consideration of thermal contact resistance is not feasible. Specific boundary condition available in Ansys (fan boundary condition) is used to split the tube as shown in post-processing (Fig. 11) to apply the characteristics curve of the diaphragm pump presenting flow rate vs pressure for greater fidelity of LM flow model. In solving the energy equation resulting temperature contours, the conjugate heat transfer is used to describe heat transfer that involves variations of temperature within solids and liquid metal due to thermal interaction between the solids and LM as seen in Fig. 12. In this simulation, surface heat is generated from mounted electronics (72 W), and heat is taken away by the AM cold plate and free convection. Heat flux post-processing report has shown that heat transferred by free convection is less than 10% and most of heat dissipated to the cold boundary condition via conduction. CFD simulations performed for in situ AM cold plates including plated ceramic tubes replicating test runs and the pre- and post-processing corresponding to in situ AM cold plate of experiment run #1 (BN-PC-IAM) are shown in Figs. 11 and 12. To check the convergence of simulation, residual curves of continuity, velocity, energy, and user-defined scalar (uds-0) are plotted. As seen in Fig. 13 after 950 iterations, the numerical model was fully converged concluding validation of postprocessing results and appropriate mesh grid size without any need for further refinement. The CFD simulation revealed key characteristics and behaviors regarding the design of the experiment including diameter of tube, thickness of block, and selection of appropriate pumping for required flow rate. The outcome of iterative simulations resulted in optimized prototypes to cool and meet the temperature limit of mounted electronics. However, verification and validation of models are required before committing to performance of developed cooling prototypes. In the following section, actual thermal test is performed for fabricated prototypes. Post-processing of simulations and experimental results are compared, and CFD simulation is with an error margin below 6% compared to test results. The comparison is demonstrated at \u201cThermal test process and experimental results\u201d section (Table 2). Fig. 16 Fabricated cold plates by in situ AM and conventional techniques (similar geometry and material for pair comparison). a In situ SLS cold plate including copper-plated BN ceramic tubes creating superior bonding by fusion of copper with aluminum alloy AM powder at thermal interfaces. b In situ SLS cold plate including silver-plated BN ceramic tubes and fusion of silver with aluminum alloy AM powder has shown highest efficiency. c Conventionally assembled cold plate including three-piece SLS aluminum alloy block, gauge 20 copper off-the-shelf plate. d Conventionally assembled cold plate including three-piece SLS aluminum alloy block, gauge 20 silver off-the-shelf plate. Three-piece assembly fastened with #4-40 SS screws torqued to 4.7 lb-in. surface flatness .003 \u201c(0.08 mm) and surface roughness 64 \u03bc-in. (0.002 mm) at top surface (interface to heat source) and bottom surface (interface to Peltier cooler) for all pieces" ] }, { "image_filename": "designv11_22_0002914_978-3-030-60330-4-Figure4.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002914_978-3-030-60330-4-Figure4.5-1.png", "caption": "Fig. 4.5 Dependence of the Poisson\u2019s ratio \u03bd1 = \u03bd<110, 11\u03040> for d/a =1 on the parameters of force a and moment b interactions: a\u2014K20 = 0.3 (curve 1), K20 = 0.6 (curve 2), K20 = 0.9 (curve 3); b\u2014K30 = 0.37 (curve 1), K30 = 0.6 (curve 2), K30 = 0.9 (curve 3)", "texts": [ "46) shows that in cases, when K20 < \u22129.69 or K20 > 0.71, \u03bd1 < 0 for any values of K30. If, then at or, where. But if \u22129.69 < K20 < 0.71, then \u03bd1 < 0 for K30 < 1 4 ( 1 \u2212 \u221a 2K20 \u2212 \u221a F ) or K30 > 1 4 ( 1 \u2212 \u221a 2K20 + \u221a F ) , where F = (10 \u2212 8 \u221a 2)K 2 20 + (8 \u2212 14 \u221a 2)K20 + 9. For example, if K20 = 0.3, then F \u2248 5.34 and K30 < \u22120.43 or K30 > 0.72. If K20 = 0.5, then F \u2248 2.77 and K30 < \u22120.34 or K30 > 0.49. Finally, if K20 = 0.6, then F \u2248 1.45 and K30 < \u22120.26 or K30 > 0.34. These results are confirmed by the shown in Fig. 4.5 dependences of \u03bd1 on the parameters of force (K30 = K3/K0) and moment (K20 = K2/K0) interactions [50]. In its turn, \u03bd2 < 0 when (K0 + K3) ( p2 + ( \u221a 2 \u2212 p)2 ) < 4 \u221a 2 ( p \u2212 1/ \u221a 2 ) K2. (4.47) 104 4 Application of the 2D Models of Media with Dense and Non-dense \u2026 Since the Poisson\u2019s ratio \u03bd2 decreases with growing of the grain size, inequality (4.47) implies the necessary condition for its negativity, which is obtained by substituting the maximum possible value p = 1 into (4.47): K3 < \u221a 2K2 \u2212 K0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002060_j.wear.2016.11.002-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002060_j.wear.2016.11.002-Figure2-1.png", "caption": "Fig. 2. A typical model of abrasive wear by a conical particle (after [12]).", "texts": [ " Therefore, to mimic actual circumstances, variants of the standard procedure must be made to obtain the type of wear information required for engineering purposes. Aside from changes in the loading weight and sliding distance, the rate of sand flow, abrasive characteristics, and test duration can be reconfigured. Research has shown that approximately 200 wheel revolutions is adequate to create a steady wear rate and that multiple shorter tests could be run instead of a single long test to protect the rubber wheel [11]. Abrasive wear can be described as a hard conical particle penetrating and sliding within a softer material as shown in Fig. 2. In a typical abrasion function, (1), the abrasive wear is quantified as a volume loss generated by a single conical particle sliding over a distance Li [12,13]. \u03c0 \u03b8 = \u22c5 \u22c5 ( )V W L H 2 tan 1i i i Eq. (1) is the typical abrasion function, where \u03c0 \u03b8( \u22c5 )2/ tan represents a wear coefficient and is dependent on the ductility of the ified rubber wheel abrasion test, Wear (2016), http://dx.doi.org/ material, interfacial shear strength, and particle shape [14]. In order to use friction energy to study abrasive wear, the friction expression \u03bcWi is introduced via Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure15-1.png", "caption": "Fig. 15. Structural analysis result of yaw bearing.", "texts": [], "surrounding_texts": [ "The reliability of the structural analysis was evaluated ex- perimentally. The stress level at the same loading condition was compared analytically and experimentally at the same position of the test rig. The applied loading was an extreme load case for a pitch bearing. A strain gauge was used to measure the stress level of the test rig. Ten strain gauges of two different types were applied at key locations in the test rig, as shown in Figs. 16-18. The gauge type was chosen according to the shape of the mounted parts, and the specifications of each type are shown in Table 3. Fig. 19 shows the stress results of the different approaches. The analytical and experimental results show good agreement in terms of the tendency and magnitude of stress. Therefore, it may be concluded that the analysis method is suitable for estimating the test rig\u2019s stress level." ] }, { "image_filename": "designv11_22_0001720_ecce.2014.6953594-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001720_ecce.2014.6953594-Figure1-1.png", "caption": "Figure 1. A schematic of the stator of an electric machine including end windings ferrofluidic enclosures", "texts": [ " The goal is to exploit the thermal and magnetic characteristics of TSFFs such that not only the fluid will circulate over the end winding of machine, but existing thermal and magnetic fields (leakage field) in this part of the machine will provide the energy needed for pumping (circulating) the cooling fluid. The cooling mechanism will therefore require no additional moving mechanical parts, pumps, or sensors. The behavior of the fluid will provide the method of pumping and cooling of the motor (thus increasing the machine potential energy density and efficiency). This approach has been implemented in the cooling of electromagnetic devices such as power transformers and electromagnetic stirrers [3,4,5,6]. In this application, the end winding of electric machine is simply immersed in a FF (Fig. 1); the strength of the magnetic field around the windings [6] is greatest in the immediate vicinity of the windings, and falls off rapidly with increasing distance from the core. Due to the magnetic field gradient, the FF is drawn toward the device. However, as the copper coils generate Ohmic losses, the temperature of the FF rises (approaches to Curie temperature of the FF) as it moves toward the windings, resulting in loss of magnetization and so in loss of attraction through windings. Then the heated FF is replaced with colder FF using the attraction exerted by the electromagnetic device on the colder FF" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002241_tie.2020.2967663-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002241_tie.2020.2967663-Figure5-1.png", "caption": "Fig. 5. (a) The angles between the magnetic axes of each phase. (b) Positive- and negative-sequence components in SRF.", "texts": [ " 1 illustrates the 2D cross-section of the LVPMM prototype with 12 slots overlaying 10 pairs of PM poles and no-load magnetic field distribution. The windings are arranged in the sequence of ABC. Because the LVPMM breaks at both ends, compared with a conventional symmetrical winding surface-mounted PM rotating machine, phase B of the LVPMM is not equivalent to phase A and phase C. It is the socalled longitudinal end effect in linear motors, which significantly affects the smoothness of thrust force. The asymmetry is represented in Fig. 5 (a), where \u03b81, \u03b82 and \u03b83 are the angles between the magnetic axes of stator winding vectors. The angles are no longer 120 electrical degrees. The self-inductance and mutual inductance waveforms and fast Fourier transform (FFT) distributions are shown in Fig. 2. The waveforms are obtained from the LVPMM finite element analysis (FEA) model with ANSYS. Maximum length of Authorized licensed use limited to: Carleton University. Downloaded on July 27,2020 at 16:28:52 UTC from IEEE Xplore. Restrictions apply", " A, zero-sequence components are neglected. Unbalanced three-phase back EMFs [\ud835\udc38\ud835\udc4e \ud835\udc38\ud835\udc4f \ud835\udc38\ud835\udc50] \ud835\udc47 or three-phase currents [\ud835\udc3c\ud835\udc4e \ud835\udc3c\ud835\udc4f \ud835\udc3c\ud835\udc50] \ud835\udc47 can be calculated according to: [ \ud835\udc45\ud835\udc4e \ud835\udc5d \ud835\udc45\ud835\udc4f \ud835\udc5d \ud835\udc45\ud835\udc50 \ud835\udc5d ] = 1 3 [ 1 \ud835\udc4e \ud835\udc4e2 \ud835\udc4e2 1 \ud835\udc4e \ud835\udc4e \ud835\udc4e2 1 ] [ \ud835\udc45\ud835\udc4e \ud835\udc45\ud835\udc4f \ud835\udc45\ud835\udc50 ] (11) [ \ud835\udc45\ud835\udc4e \ud835\udc5b \ud835\udc45\ud835\udc4f \ud835\udc5b \ud835\udc45\ud835\udc50 \ud835\udc5b ] = 1 3 [ 1 \ud835\udc4e2 \ud835\udc4e \ud835\udc4e 1 \ud835\udc4e2 \ud835\udc4e2 \ud835\udc4e 1 ] [ \ud835\udc45\ud835\udc4e \ud835\udc45\ud835\udc4f \ud835\udc45\ud835\udc50 ] (12) where \ud835\udc4e = \ud835\udc52\ud835\udc57 2\ud835\udf0b 3 or (\u2212 1 2 + \ud835\udc57 \u221a3 2 ) , \ud835\udc4e2 = \ud835\udc52\u2212\ud835\udc57 2\ud835\udf0b 3 or (\u2212 1 2 \u2212 \ud835\udc57 \u221a3 2 ), \ud835\udc45 represent \ud835\udc38 or \ud835\udc3c. The superscript \ud835\udc5d represents the positive-sequence component and \ud835\udc5b represents the negativesequence component. Fig. 5 (b) shows the relationship between the positive- and negative-sequence components in d- and q-axis SRF and stationary \u03b1\u03b2 reference frame. The positive- and negativesequence components can be expressed as: \ud835\udc45\ud835\udc51\ud835\udc5e \ud835\udc5d = \ud835\udc52\ud835\udc57\ud835\udf14\ud835\udc61\ud835\udc45\ud835\udefc\ud835\udefd \ud835\udc5d (15) \ud835\udc45\ud835\udc51\ud835\udc5e \ud835\udc5b = \ud835\udc52\u2212\ud835\udc57\ud835\udf14\ud835\udc61\ud835\udc45\ud835\udefc\ud835\udefd \ud835\udc5b (16) where \ud835\udc45\ud835\udc51\ud835\udc5e \ud835\udc5d = \ud835\udc45\ud835\udc51 \ud835\udc5d + \ud835\udc57\ud835\udc45\ud835\udc5e \ud835\udc5d , \ud835\udc45\ud835\udc51\ud835\udc5e \ud835\udc5b = \ud835\udc45\ud835\udc51 \ud835\udc5b + \ud835\udc57\ud835\udc45\ud835\udc5e \ud835\udc5b , \ud835\udc45\ud835\udefc\ud835\udefd \ud835\udc5d = \ud835\udc45\ud835\udefc \ud835\udc5d + \ud835\udc57\ud835\udc45\ud835\udefd \ud835\udc5d, \ud835\udc45\ud835\udefc\ud835\udefd \ud835\udc5b = \ud835\udc45\ud835\udefc \ud835\udc5b + \ud835\udc57\ud835\udc45\ud835\udefd \ud835\udc5b. The first term \ud835\udc52\ud835\udc57\ud835\udf14\ud835\udc61\ud835\udc45\ud835\udefc\ud835\udefd \ud835\udc5d, rotating counterclockwise, denotes the positive sequence, and the second term \ud835\udc52\u2212\ud835\udc57\ud835\udf14\ud835\udc61\ud835\udc45\ud835\udefc\ud835\udefd \ud835\udc5b , rotating clockwise, denotes the negative sequence" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003847_j.mset.2021.08.005-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003847_j.mset.2021.08.005-Figure1-1.png", "caption": "Fig. 1. PhysicalModel showing assemblywith channel angle 90 and corner angle 20.", "texts": [ " It is a FEM-based software used for studying different machining and metal forming processes that help in observing an efficient graphical interface for convenient analysis and preparation of data [19\u201335]. The substance of the workpiece considered in the simulation is Ti-6Al-4V alloy. Table 1 shown above shows the composition, physical and mechanical properties of the workpiece material. The workpiece is passed through the channel in the shape of a billet with a diameter of 20 mm and a length of 70 mm as shown in Fig. 1 below. The physical model is made using ANSYS software. The process was considered to be taking place at 22 C. The workpiece consists of 44,255 nodes and 9878 elements which were used for domain discretion and meshing. The last requirement was to set up the friction coefficient between the die and workpiece material, constraints, and friction. The tolerance between the die and the workpiece will be automatically achieved by ANSYS. Table 2 below shows the various fixed parameters that were taken during the simulation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure10.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure10.12-1.png", "caption": "Figure 10.12 (a) Balancing arbor and (b) stand", "texts": [ " The teeth are formed by the tiny D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 10 Surface grinding 10 150 face of the wheel flange is an annular dovetail groove which holds balance masses, usually two. These masses can be locked in any position round the groove. When the wheel is mounted correctly and the periphery has been dressed, the collet assembly is removed from the spindle. The balance masses are then removed. A balancing arbor is inserted in the bore. The balancing arbor has a taper identical to that on the spindle and has equal size parallel diameters at each end, Fig.\u00a010.12(a). This assembly is placed on a balancing stand, Fig.\u00a010.12(b), which has previously been set level. The wheel is allowed to roll on the knife edges and is left until it comes to rest, which it will do with the heaviest portion at the bottom. A chalk mark is made at the top of the wheel, opposite the heaviest portion. The masses are then replaced in a position opposite to each other and at right angles to the chalk mark. The masses can now be moved equally a little way towards the dependent upon the size of wheel, e.g. a 250 mm diameter by 25 mm wide wheel would typically require a 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000257_j.jfranklin.2019.06.004-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000257_j.jfranklin.2019.06.004-Figure1-1.png", "caption": "Fig. 1. side view of suspended floater system.", "texts": [ " (25) : \u2212\u03bc Y i \u2212I ] < 0, \u22122\u03c1N 1 + \u03c12 I + \u03baI < 0 (25) here \u03c1 is a predefined scalar. emark 3. The stability conditions in [37] are not numerically solvable, the H \u221e index \u03b3 ust be given at first, then the stability conditions can be solved, so certain conservatism xists. In Theorem 2 , the stability and robustness conditions are in convex form, and \u03b3 can e minimized directly by convex optimization method. . Performance analysis The braking control task of suspended floater following control system is adopted here o verify the effectiveness of the proposed method. As shown in Fig. 1 , the kinetics and ynamics of the system can be described as: \u02d9 = f 1 (x, u) + f 2 (x, d ) z = C 1 x = C 2 x \u03b8 is the swing angle of the suspended rope; v is the velocity of the motion platform; u is the control force of the platform motor; and d is the jet propulsion force from the floater which can be seen as disturbance. Parameter m is the mass of the floater; M is the mass of the motion platform; l is the rope length; g is the gravitational acceleration; C r is the friction damping coefficient between motion platform and guide rail; C s is the friction damping coefficient of the connection point between rope and platform" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002743_ccc50068.2020.9188925-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002743_ccc50068.2020.9188925-Figure2-1.png", "caption": "Fig. 2: Demonstration of a fixed-wing UAV in the global coordinate frame.", "texts": [ " When three above assumptions about the nominal formation hold, there exist p (t) \u2261 p\u2217 (t) and p\u2217f (t) = \u2212\u03a9\u0304\u22121 ff \u03a9\u0304flp \u2217 (t), and this problem can be transformed into control the followers\u2019 positions pf (t) \u2192 p\u2217f (t) as t \u2192 \u221e. 3 Main Results A fixed-wing UAV\u2019s planar movement can be simplified into the unicycle model with Dubins constraints, which is suitable for high-level path planners. The model of the ith UAV can be expressed as x\u0307i = \u03bdi cos\u03d1i, y\u0307i = \u03bdi sin\u03d1i, \u03d1\u0307i = i, (5) where xi, yi \u2208 R, \u03bdi \u2208 R, i \u2208 R, and \u03d1i \u2208 R represent the planar coordinates, linear speed, angular speed, and the yaw angle of the ith UAV respectively. In the global coordinate, a demonstration of fixed-wing UAV is given in Fig. 2. Meanwhile, the Dubins constraints of the ith fixed-wing UAV are defined by 0 < \u03bdmin \u2264 \u03bdi \u2264 \u03bdmax, \u2212 max \u2264 i \u2264 max, (6) where \u03bdmin and \u03bdmax are the positive minimum and maximum linear velocities, and max is the positive maximum angular velocity. The minimum turn radius of all UAVs is defined as rmin = \u03bdmin/ max. For the purpose of designing the controller, the yaw vector i = [cos\u03d1i, sin\u03d1i] T \u2208 R 2 and its perpendicular vector \u22a5 i = [\u2212 sin\u03d1i, cos\u03d1i] T \u2208 R 2 are introduced. Since pi = [xi, yi] T and \u0307i = \u22a5 i \u03d1\u0307i, (5) is modified as p\u0307i = i\u03bdi, \u0307i = \u22a5 i i", " (9) Assume that the leaders fly with the same speed and the same yaw angle simultaneously, and the directional unit vector is qi \u2208 R 2, i \u2208 V and the positive magnitude of the speeds is c \u2208 R for all the leaders. Thus p\u2217 (t) = p\u2217 (0) + tcq , h (t) = q , (10) where q = [qT1 , q T 2 , \u00b7 \u00b7 \u00b7 , qTn ]T . If cqi, i \u2208 V can be estimated by the followers in finite time [12] and no Dubins constraints of the followers are considered, then the ith follower\u2019s tracking control scheme, i \u2208 Vf , can be designed as \u03bdi = T i \u23a1 \u23a3\u2212 \u2211 j\u2208Ni \u03c9ij(pi \u2212 pj) + cqi \u23a4 \u23a6 , i = ( \u22a5i ) T \u23a1 \u23a3\u2212 \u2211 j\u2208Ni \u03c9ij(pi \u2212 pj) + cqi \u23a4 \u23a6 , (11) where the actual meaning of this control scheme is shown in Fig. 2. Note that this controller is activated within each UAV\u2019s local coordinate frame, and only relative position and desired yaw angle measurements are required. The rotational transformation form a global coordinate to a UAV\u2019s local coordinate does not affect the structure of the controller. It can be viewed a barycentric-coordinate based control approach. In the rest part of this paper, the proposed control strategy can be executed in the local coordinate frame as well. Theorem 3. When Assumptions 1-3 hold, if the leaders maintain (10), then the tracking error \u03b4pf of the followers can converge globally to zero under the control scheme (11)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003339_1464420721990049-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003339_1464420721990049-Figure12-1.png", "caption": "Figure 12. Melt pool track temperature evolution during the model\u2019s element birth; thermal history of a probed point in (a) layer \u00bc1, (b) layer \u00bc32nd (c) layer \u00bc70th, and (d) layer\u00bc 140th.", "texts": [ " The resolution of the FEM model was selected to be high enough to guarantee stabilization as well as an accurate cooling rate but keeping an affordable computational time. Thus, the Courant Friedrich Lewy condition27 of the equation (8) must be fulfilled C \u00bc uDt Dx Cmax (8) where C is the Courant number, u \u00bc 14:167mm=s is the velocity of the scanning laser, Dx \u00bc 35mm is the cube length, and Cmax \u00bc 1 for the direct solver. Then, the time step selected for solving the transient heat transfer phenomena is Dt \u00bc 0:2 s which is low enough to capture the temperature evolution along the nodes. Figure 12 shows the temperature contour during first up to the 140th track deposition. The measured running time is summited to the end of the process solution equal to 123 h. The simulations run in parallel computer using a total number of 24 CPUs. Dependence of pore descriptors on solidification cooling rate. The aim of thermal localized study on cubic components is a step into gaining insight into how porosity microstructure and solidification cooling are linked. The size, shape, and distribution of the individual aligned interlayer pores in the cubic component led to proper characterization of the relationship between the DED process parameters and mechanical behavior" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001323_ihmsc.2014.92-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001323_ihmsc.2014.92-Figure6-1.png", "caption": "Figure 6. The model of robot and the environment of simulation.", "texts": [ " In order to have research on the self-organized behaviors in swarm robotics, the robot named swarm-bot is developing. Relative positioning systems of robots based on infrared technology, using received signal strength indicator measurements is to calculate the relative range and bearing[14]. The experiment is implemented in Webots. Before the simulate experiments, the reasonable assumptions and constraints are necessary. In the aspect of hardware, the paper builds a closed twodimensional circular arena in Webots (Fig.6). In each robot, eight infrared distance sensors are distributed uniformly, which can perceive the boundary of the arena. Supposing to ignore the time of acceleration and deceleration of motor and angle error of the robot, there are 16 LED infrared transmitters and 8 infrared receivers on the top of the robot which is used to acquire the local positioning of the robots around. The robot also has a comprehensive two-way radio frequency antenna. In the aspect of software, every robot has a clock which can show the current time of system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002749_s11071-020-05932-9-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002749_s11071-020-05932-9-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of a 3D pendulum", "texts": [ " k k denotes the 2-norm of a vector or the induced norm of a matrix. \u00f0 \u00de denotes the 3 3 skew-symmetric matrix of a vector. sgn \u00f0 \u00de denotes the signum function. For a vector x \u00bc x1; x2; . . .; xn\u00bd T2 Rn and a positive scalar c, xc \u00bc xc1; x c 2; . . .; x c n T and sigc x\u00f0 \u00de \u00bc x1j jcsgn x1\u00f0 \u00de;\u00bd x2j jcsgn x2\u00f0 \u00de; . . .; xnj jcsgn xn\u00f0 \u00de T. 2.1 Attitude kinematics and dynamics of the 3D pendulum Suppose that the 3D pendulum consists of a rigid body connected by a fixed and frictionless pivot subjected to a constant gravitational force as shown in Fig. 1. The 3D pendulum is fully actuated by three-axis control torques and is allowed to rotation around the pivot with three degrees of freedom. Two coordinate frames are introduced to describe the attitude motion of the 3D pendulum. They are the inertial reference frame which the origin is located at the pivot, the first two axes lie in the horizontal plane, and the third axis is vertically along the direction of gravity, and the bodyfixed frame which the origin is fixed at the pivot and the three axes coincide with the principal axes of inertia" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003340_tia.2021.3058200-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003340_tia.2021.3058200-Figure7-1.png", "caption": "Fig. 7 Equipotential and flux density distributions due to PM excitation. (a) 18/17-pole IFC-BFPMM. (b) 18/13-pole IFC-DSPMM.", "texts": [ " 6 compares electromagnetic torque of the initial and optimized designs of both machines. It can be observed that the electromagnetic torques of the optimized 18/17-pole IFCBFPMM and 18/13-pole IFC-DSPMM significantly boost. To get further insight into the characteristics of the optimized two machines, such performances as open-circuit field distribution, open-circuit phase flux linkage, phase back-EMF, dq-axis inductances, torque capability, efficiency and power factor, cogging torque, unbalanced magnetic force (UMF), and demagnetization of PMs are minutely discussed here. Fig. 7 shows the equipotential and flux distributions due to PM excitation. It can be observed that each PM in both machines has flux leakage through the machine outer region or directly through the air gap. Due to the larger number of PM pieces, the 18/17-pole IFC-BFPMM has more flux leakage than the 18/13-pole IFC-DSPMM. Fig. 8 compares the open-circuit phase flux linkages between two machines with the same number of turns per coil. Both machines have the bipolar and sinusoidal phase flux linkages since the coils differing by 180 elec" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003676_s00170-021-07413-8-Figure28-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003676_s00170-021-07413-8-Figure28-1.png", "caption": "Fig. 28 Material effect on wrinkling\u2019s defect", "texts": [ " So, the variation of compressive stress due to change in the cavity pressure has a significant effect on wrinkle waves. To study the influence of the material on the wrinkling phenomenon in HDDRP with inward flowing liquid, the pure copper sheet with a thickness of 1 mm and the geometry represented in Fig. 4 was used. The copper sheet is assumed to be isotropic as mentioned in Sect. 2. Both simulations were done under the same conditions for the accurate comparison. The final shape of the simulation is demonstrated in Fig. 28. Based on this figure, when the copper blank was used, the amplitude of the wrinkle wave will be formed at a lower value (Fig. 29). However, at the same time, the number of wrinkling waves increases. That is why the wave height is reduced. In other words, along the circumference of the cone, more waves with less height can be seen. However, the wrinkling height in the steel cone is greater because wrinkling in materials with higher yield stress is more severe in comparison to the materials with lower yield stress [44]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001279_s0263574714000149-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001279_s0263574714000149-Figure2-1.png", "caption": "Fig. 2. The multi-body system model.", "texts": [ " (1) In (1) S is the motion dimension of the space (S = 6 for the spatial mechanisms and S = 3 for planar mechanisms), n is the number of mechanism mobile bodies and Ci the number of joints of class i (the class of a joint is given by the number of constraints introduced). In fact this is the structural model that is fundamental to the category identified by Gogu2 and Mueller4 as Chebychev\u2013Kutzbach\u2013Gru\u0308bler (CKG) formulae. As each body is characterized by S independent degrees of freedom, the instantaneous position of a mechanism with n bodies is thus characterized by S\u00b7n generalized coordinates. This is illustrated in Fig. 2 for a planar mechanism for which the position is fully described by S\u00b7n = 3\u00b75 = 15 generalized coordinates. To obtain the kinematic equations of this system, the geometrical constraints introduced by the joints are expressed by algebraic equations. In total, the number of equations introduced by all joints is i\u00b7Ci . The number of independent generalized coordinates, i.e. those that cannot be calculated from the constraint equations, equals the mechanism mobility Nqi = M = 6 \u00b7 n \u2212 \u2211 i \u00b7 Ci. (2) Therefore the formulation of the position kinematics for the mechanism modeled as a multibody system must include M equations for the driving motions, in addition to the i \u00b7 Ci constraint equations, reaching a total number of M + i \u00b7 Ci equations. http://journals.cambridge.org Downloaded: 18 Mar 2015 IP address: 128.233.210.97 As an illustration, the planar mechanism given in Fig. 2 has 5 bodies and 7 joints, each of the latter introducing two geometric constraints. The mobility is then M = 3\u00b75\u20137\u00b72 = 1. The mobility formula shows thus that a number of 7\u00b72 = 14 constraint equations can be formulated for the kinematic model (two for each constraint) expressing the algebraic conditions for the revolute joints A. . .F and the translational joint G, respectively. A further equation expressing the driving motion is needed in order to completely define the mechanism in case it is kinematically driven", " S stands for spatiality with S = 6 for spatial mechanisms and S = 3 for the planar ones. As the number of unknowns\u2014the generalized accelerations and Lagrange multipliers\u2014exceeds the number of equations, one needs to consider the kinematic equations as well for the purpose of numerical integration. The differential algebraic system (DAE) with the typical general form { \u0308(q, t) = 0 mq\u0308 \u2212 J T \u03bb = Qex (4) has thus in total S \u00b7 n + i \u00b7 Ci equations (S\u00b7n differential and i \u00b7 Ci algebraic equations). For example, for the planar mechanism in Fig. 2, each body provides 3 equilibrium equations and, for all five bodies, 3\u00b75 = 15 dynamic equations should be written. Thus, the constraints and dynamic equations constitute a system of 6\u00b7n + i\u00b7Ci = 3\u00b75+7\u00b72 = 29 differential-algebraic equations (DAE). Note that the Chebyshev-Kutzbach-Griibler (CKG) mobility formula is characteristic only for mechanisms modeled as multi-body systems (MBS). Other than the number of bodies and the type and number of joints, it offers relevant information on the kinematic formulations, such as the number and structure of kinematic equations, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003454_j.optlastec.2021.107058-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003454_j.optlastec.2021.107058-Figure3-1.png", "caption": "Fig. 3. (a) Parallel direction samples (PD); (b) vertical direction samples (VD); (c) the profile of the tensile test bar.", "texts": [ " The precipitated phases in the sample were analyzed using JEM-2100 field emission transmission electron microscope (TEM). Before the metallographic observation, the samples were ground, polished and etched successively. The etching solution is HNO3 and C2H5OH with a volume ratio of 1: 5, and the sample\u2019s surface is immersed in the solution for 40 s. The microhardness tests were measured by the HV-2000 digital microhardness tester with a load of 200 g and a dwell time of 10 s. The measurement is carried out from deposition layer in x-z plane (Fig. 3a) to the heat affected zone (HAZ) in the substrate at an interval of 500 \u03bcm. Tensile samples were obtained from two directions. The samples parallel to the laser scanning direction are identified as parallel direction (PD), and the vertical scanning direction are identified as vertical direction (VD). As shown in Fig. 3, the specimens were cut according to the Chinese national standard GB/T 228.1-2010. A high sensitive tensile machine, MTS E45.105, was used for the test at room temperature, and the strain rate is 0.001 s\u2212 1. The yield strength, ultimate tensile strength and fracture strain were obtained as direct output from the tensile testing machine. Three independent specimens were tested and the average values were obtained to ensure the reliability. In order to Fig. 2. Photographs of multilayer sample produced by LMD: (a) top view, (b) side view", " Table 2 Process details parameters of LMD. Number of layers 1, 2 3, 4, \u2026 ,20 Laser power (W) 1000 1000 Scanning speed (mm/s) 6 10 Layer thickness (mm) 0.8 1 Overlap ratio 50% 50% Powder feeding rate (g/min) 10.5 13 J. Zhang et al. Optics and Laser Technology 140 (2021) 107058 investigate the fracture mechanism, the fracture morphology was observed by SEM and TEM. To research the tensile property of LMD-processed samples, the quasi-static tensile tests were carried out in two directions, as shown in Fig. 3. The yield strength, tensile strength and elongation at room temperature are listed in Table 3. For the PD direction (Fig. 3a), the average tensile strength and yield strength of the tensile specimen are 715 MPa and 530 MPa, respectively. For the VD direction (Fig. 3b), the average tensile strength and yield strength are 760 MPa and 532 MPa, respectively. Comparing with the American Aerospace Materials Technical Specification AMS5754, the tensile strength and yield strength of LMD specimens in both tensile direction are higher than the stipulated 690 MPa (in tensile strength) and not less than 275 MPa (in yield strength) at room temperature. The yield strength, tensile strength of LMD samples are about 1.9 times the standard value, and its tensile strength is 4%-10% higher than the standard, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000485_etfa.2019.8869079-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000485_etfa.2019.8869079-Figure4-1.png", "caption": "Fig. 4. CAD part (left) and the generated path using MAT strategy (right).", "texts": [ " MAT was first introduced by Blum to describe shapes with medial axis defined as \u201cloci of centers\u201d of locally maximal spheres inside an object [23]. In two dimensions (2D) the MAT would be the \u201cloci of centers\u201d of locally maximal circles inside the region of a 2D shape, Fig. 3. Recently, it has been identified as a solution for MAM path planning. The path planning strategy starts by filling the part from inside towards outside. By using this strategy, it is possible to reduce imperfections like pores and gaps related to MAM. Fig. 4 shows a path example applied to a part based on MAT strategy for DED. Our proposed approach passes by importing to an AutomationML file the data from our path generator combined with the process parameters. At the path planning phase, once the path is created, Fig. 4 (right), we can generate a numerical control (NC) file containing all the points of the path. The points are extracted from the NC file and stored into an commaseparated values (CSV) file, and later in an AutomationML file by means of the AutomationML Generator. At the same time, when the points are uploaded to the AutomationML file, the data are stored into a CAEX tree structure. Fig. 5 illustrates the architecture of the AutomationML generator. Starting from the CSV file, which holds the path information from the NC file generated form the CAM software" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000439_j.mechmachtheory.2019.103612-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000439_j.mechmachtheory.2019.103612-Figure2-1.png", "caption": "Fig. 2. (a) The constant-force bench press and (b) detail of the resistance system based in a CFM.", "texts": [ " Additionally, the components design and selection have been made to minimize the moving masses of the mechanism, and therefore to reduce the influence of the inertial forces, as well as to reduce the friction forces in the contact between the rollers and the cam. Although commercial Smith machines may work with higher resistance capacity and greater bar displacement, it should be noted that the proposed constant-resistance bench press is primarily oriented to this research work and therefore the above design requirements are considered suitable for its evaluation in future studies with subjects. Fig. 2 shows the 3D CAD model of the constant-resistance bench press, which includes one CFM attached to each of the two ends of the barbell. Each CFM has a stroke of 0.7 m, which allows the same displacement of the barbell. It can be mounted up to 28 springs in pairs to maintain the stability of the roller axle (one on each side of the roller), where each pair of springs increases the resistance force by 5 kp, with a maximum resistance of 70 kp when all the spring are mounted. Since there is one CFM at each barbell end, the bench press has a maximum resistance of 140 kp" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002950_tbme.2020.3042115-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002950_tbme.2020.3042115-Figure5-1.png", "caption": "Fig. 5. Diagram of experimental setup demonstrating the arrangement of robotic guidewire and porcine artery during data acquisition.", "texts": [ " After each in-plane acquisition, the 3-axis motion stage was used to translate the guidewire in the elevation direction to allow acquisition of 3D datasets, where each plane of data was acquired using the robotically steerable guidewire and adjacent planes were acquired by stepping the motion stage out of plane with a step size of 0.2 mm. Finally, the same 3D acquisition scheme was used to image an ex vivo porcine carotid artery with an inner diameter of \u223c4 mm at a depth of 5 mm. The artery was submerged in saline solution and placed with the patent (unobstructed) lumen in front of the system to mimic the clinical use of the system (Fig. 5). A custom acrylic stand was used as a frame, and sutures located deep to the imaging field of view were used to hold the artery in place, preventing motion and maintaining form as in [73], [74]. Examination of the magnitude and phase of the electrical impedance of the completed device with air loading (Fig. 6) Authorized licensed use limited to: San Francisco State Univ. Downloaded on June 19,2021 at 06:41:28 UTC from IEEE Xplore. Restrictions apply. indicates a series resonance of 17 MHz and electrical impedance near 50 \u03a9" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001229_tmag.2014.2360232-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001229_tmag.2014.2360232-Figure4-1.png", "caption": "Fig. 4. One-turn inductor: various volume corrections (field lines and Joule power density). Left to right: from Biot\u2013Savart solution (SP 1) to volume corrections to circular (radius 5 mm), square and rectangular conductors (constant surface area) (SPs 2, frequency f = 1600 Hz).", "texts": [], "surrounding_texts": [ "At any step of the process, the addition of a region (e.g., a magnetic and/or conducting core) defines an SP p for calculating the associated reaction field [3]. The required VSs (2a-b) need the calculation of the already present fields as sources, both Biot\u2013Savart fields and their possible volume corrections (all gathered for the current solution q). D. Inductance and Resistance Calculation The self-inductance of a wire conductor, and the possible mutual inductances with other wire conductors, can be calculated via double integral Neumann formulas [7], [8]. The resistance can be approximated as well. After a volume correction SP p, the corrected inductance is advantageously obtained with the solution ap only in s,p and c,p , i.e., via s,p = (js,p, ap) s,p , c,p = (jp, ap) c,p (13a-b) defining the total linkage magnetic fluxes s,p and c,p [4], thus as the new global value without any reference to the wire inductance approximation. This is a valuable key feature of the proposed method. An added region in an SP gives an inductance change that is calculated by a volume integral limited to the added region (via the reciprocity theorem [6]), which is another key advantage of the SPM." ] }, { "image_filename": "designv11_22_0000923_icpe.2015.7167962-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000923_icpe.2015.7167962-Figure1-1.png", "caption": "Fig. 1. Reference frames of PMSM.", "texts": [ " The stator resistance estimation is described in [7] and [10], and also online estimation methods using the signal injection are suggested in [8], [9], and [11]. From the view point of turbo compressor applications, this paper analyzes the control performance of the current model-based algorithm with respect to motor parameter variations in normal operating speed region and then suggests the adaptive sensorless control method by adopting the back emf constant estimator, which does not depend on the signal injection. The Matlab/Simulinkbased simulation and experimental results are provided to prove the validity of the proposed adaptive algorithm. Fig. 1 shows reference frames of PMSM. The - and - reference frames represent the stationary and the synchronous reference frames, respectively. The -axis coincides with the north pole of the permanent magnet rotor. The d-q reference frame is an estimated synchronous reference frame used in the sensorless vector control. and are the actual and the estimated rotor positions, respectively. The angular position error is defined as (1). (1) The voltage and torque equations of a non-salient PMSM in the estimated synchronous reference frame are Adaptive Sensorless Control of High Speed PMSM with Back EMF Constant Variation Jin-Woo Lee Dept" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002274_j.mechmachtheory.2019.103748-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002274_j.mechmachtheory.2019.103748-Figure8-1.png", "caption": "Fig. 8. 3-UPU manipulator \u2013 schematic representation, the couplings are represented by labels A to L.", "texts": [ " Experimental data can be used with Eqs. (26) , (27), (32) and (33) in which case numerical integration may be needed. This section is dedicated to the friction analysis of a complex machine with parallel kinematic chain aiming to establish its efficiency. The machine selected as an example is a parallel robot called 3-UPU. The 3-UPU [27] is a three-degrees-of-freedom spatial parallel manipulator composed of three limbs connected to a base and a platform by means of universal or Hooke joints as shown in Fig. 8 . The labels are related to Figs. 9 and 10 and explained further on. Each limb has one actuated prismatic joint connected to two universal joints. The limbs are assumed to be identical and their connections to the base and platform are symmetrically distributed (the centres of the universal joints form two equilateral triangles, one on the base and the another on the platform, parallel to the triangles shown in Fig. 8 ). The proposed method can be applied to asymmetrical machines in general, but symmetry is convenient because it allows a more concise description. Moreover, the system is assumed to be quasi-static and, accordingly, the inertia effects and control errors are neglected. Evaluation of the overall efficiency of the 3-UPU requires the computation of motions and actions during the execution of a task. Normally, the task is prescribed by the desired movement of the platform carrying a payload. Since the 3-UPU is only capable of translation, the description of the task is equivalent to the description of the motion of any point on the platform" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001720_ecce.2014.6953594-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001720_ecce.2014.6953594-Figure8-1.png", "caption": "Figure 8. End winding temprature distribution, (A) TSFF cooled (velocity vectores are shown as well), (B) air cooled (h=50 W/m2-K)", "texts": [ " This function can be used for wide ranges of magnetic field intensity and temperature: M=Ms L(\u03be), (20) L(\u03be)=coth\u03be-1/\u03be (21) \u03be= \u03bc0mH/kBT (22) Here, Ms is the saturation magnetization of the magnetic fluid, and L(\u03be) represents the Langevin function. The Langevin parameter \u03be is the ratio of magnetic to thermal energies [12-15]. In this parameter m and kB denote the magnetic moment of magnetic particles in the ferrofluid, and Boltzmann constant, respectively. Temperature results for two cases of cooling the end winding surfaces are presented in Fig. 8. Fig. 8(A) considers thermomagnetic convection effects of a TSFF filled enclosure while Fig. 8(B) considers typical air cooling (from wafters on rotor shaft with equivalent convective heat transfer coefficient as 50 W/m2-K [7]). As Fig. 8(A) illustrates (temperature profile and flow velocity vectors), temperature sensitive magnetic fluid motion occurs due to both the gradients of the magnetic field and the temperature. The magnetic fluid is attracted toward regions with larger field strength, while near the heat source the fluid temperature approaches the Curie temperature of the ferrofluid. In this region the fluid loses its attraction to the magnetic field, and is displaced by colder fluid [12, 15]. This application utilizes FFs whose magnetic properties are strongly influenced by temperature" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure13-1.png", "caption": "Fig. 13. Initial rigid rotation and displacement in the assembled test setup for axial node.", "texts": [ " Based on the predicted maximum loads that the structural node can sustain in different loading conditions, the loads required in the experimental tests can be calculated, considering the deterministic structural system of the test rig. Details of the applied loads and the dimensions of test rig in different tests are shown in Fig. 12. In the experimental test of the axial node, the test rig is firstly assembled according to the configuration shown in Fig. 12 and then the node is placed in the designed location. It shall be mentioned that initial displacements and rotations of the rigid rotating plates (purple plates) are observed after setting up the test rig, as shown in Fig. 13. This is because of the difference between the diameters of bolt and its hole, leading to a movement of the bolt in the direction of the force being transferred by that bolt. The accumulation of these bolt movements may result in the rigid displacements and rotations in the rotating plates, and hence change the force condition from what the mechanism is designed for. This defect is considered in the design of a generalised test setup in Section 3. Due to this initial movement, numerical simulation is conducted again, and the predicted results are found to be very different" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001199_rev.2014.6784240-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001199_rev.2014.6784240-Figure4-1.png", "caption": "Figure 4. Drive unit panel panel for remote workplace Load elevator", "texts": [], "surrounding_texts": [ "Establishment of the network of the remote workplaces was triggered by the needs of HFTM, Switzerland. It was developed in cooperation with Siemens Schweiz AG. Initial goal of the development was to give regular students more opportunities and flexibility in practical work by offering some remote facilities. However after first positive experience the network came in use also at the industrial partner for the education of professionals from industry. Building of load elevator workplace was an additional impetus for further development of the network, which now offers a number of services including: management of resources, management of users, management of access, management of policies (i.e. access policy). 978-1-4799-2024-2/14/$31.00 \u00a92014 IEEE Polytechnic of Porto (ISEP) in Porto, Portugal from 26-28 February 2014 2014 11th International Conference on Remote Engineering and Virtual Instrumentation (REV) Page 141 The network today consists of many workplaces that are in different locations and are connected to a main gateway via a proxy server, [10]. All components of each remote workplace are connected to the network. Therefore also different devices, cameras, PCs, industrial computers, programming terminals and similar, can be combined into new remote workplaces to fulfil one\u2019s current needs. The design of the transmissions of data is based on high standard protocols such as RDP, RTSP, H264, MPEG4 Codec 2. This allows a transmission with low bandwidth. Remote workplaces connected to CEyeClon can be operated easily also by using a 3G Network. In order for the user to access workplace or a service, an access session with the CEyeClon network has to be established. For this a CEyeClon viewer, a stand-alone desktop-based client, is used. The viewer is available for download at web page [1]. Its main characteristics are as follows: Cross-platform. The CEyeClon viewer can be used in any Windows releases such as Windows XP Windows 7 and Windows 8. It works also on MAC OS with Windows boot. Security. The CEyeClon Network uses the upper Ports to connect client and remote workplace. A dynamic firewall script increases the security level of the CEyeClon Network. Because of clear definite port range the system administrators can easily control remote activities of the clients. Installation. Beside the CEyeClon viewer no other installation in the client side is required. Booking system. Booking system is available for teachers/administrator. Multilingual environment. Viewer is available in English, German and French language. The CEyeClon is today rapidly growing integrated service realtime remote network used both in industry and education. Its use has increased in last year for almost 300%. III. E-TRAINING: ENERGY EFFICIENT DRIVE TECHNOLOGIES Developed training (e-materials and remote workplace) presents concept of energy efficiency by using energy efficient drives, which can, by operating in generator mode, recuperate braking energy. The learners should at the end have the capability to develop and realize software for automated machines and systems. Required preliminary knowledge and skills for enrolment in the course are basic knowledge of electric drives and control theory and an ability to analyse schemes and diagrams of industrial technical processes. The learner doesn\u2019t need to know the engineering software from SIEMENS (STEP7, STARTER) in advance since this knowledge can be also obtained during the training. The training was tested and is now used for regular education and for education in the industry (as described in the continuation). 978-1-4799-2024-2/14/$31.00 \u00a92014 IEEE Polytechnic of Porto (ISEP) in Porto, Portugal from 26-28 February 2014 2014 11th International Conference on Remote Engineering and Virtual Instrumentation (REV) Page 142" ] }, { "image_filename": "designv11_22_0003783_s40192-021-00222-7-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003783_s40192-021-00222-7-Figure1-1.png", "caption": "Fig. 1 Three-dimensional finite element model of SLM", "texts": [ " How the precipitate works in the material strengthening in the SLM AM process and the following post-heat treatment is very key for controlling and improvement of the SLM AM component. So, the PM model is combined with the FE model of SLM AM to study the precipitate contributions to material strengthening in AM process and the following post-heat treatment. Post-heat treatment can be then optimized for the improvement of the AM component. A 3D finite element (FE) model was developed based on ABAQUS software as shown in Fig.\u00a01. The model contains two regions: an AA6061 powder bed which interact with the laser beam and an AA6061 substrate. Linear fullintegration eight-node brick elements are used in finite element model. The dimensions of elements in powder bed are 35 \u00d7 35 \u00d7 25\u00a0\u03bcm, and the total number of elements and nodes is 47,520 and 52,059 respectively. The dimensions of powder bed and laser scanning region are 2.31 \u00d7 1.26 \u00d7 0.3\u00a0mm (three layers) and 1.4 \u00d7 0.7\u00a0mm. The dimensions of substrate are 4.31 \u00d7 3.26 \u00d7 0.8\u00a0mm", " The comparisons between the measured and predicted response of 6xxx series aluminum alloy to reheating in salt baths are shown in Fig.\u00a06. The comparison shows that the 1 3 precipitate evolution model is well supported by the experimental data. The material experiences reheating or remelting in the SLM AM process. In the subsequent solidification, the hardness of material increases with the generation of precipitates. The variation of hardness of Point C (the position of Point C was described in Fig.\u00a01) during the cooling stage was calculated by PE model, and the result is shown in Fig.\u00a07. From 0 to 5\u00a0s, the newly generated precipitates are less due to the higher temperature and the solid solution strengthening is the primary strengthening mechanism. With the continuous decrease in temperature, precipitates are constantly generated from the matrix, which leads to the increase in hardness. When the temperature is lower than the minimum temperature required to generate the precipitates, the hardness keeps constant", " [41] and Simar et\u00a0al. [42] shows the assumption of the homogeneous composition distribution can lead to results which can agree well with the experimental data. So, the composition is assumed to be homogeneous in current work. Fig. 6 Comparison between experimental results and numerical results: a particle number density and mean radius, b hardness 1 3 Figure\u00a08 shows the temperature distribution and the temperature history when the second layer is scanned by laser, and the position of the five points is shown in Fig.\u00a01. The laser passes through each center point of the tracks in turn, and the temperatures of the selected points show similar fluctuations. The maximum temperature of point A, point B, point C, point D and point E is 1667.3\u00a0\u00b0C, 1796.2\u00a0\u00b0C, 1834.5\u00a0\u00b0C, 1848.4\u00a0\u00b0C and 1855.6\u00a0\u00b0C, respectively. The temperature fluctuation decreases when the laser moves away. The temperature cools down rapidly as the laser moves away from the point. Due to the accumulation of heat, the maximum temperature from point A to point E increases" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001875_978-4-431-54595-8-Figure9.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001875_978-4-431-54595-8-Figure9.7-1.png", "caption": "Fig. 9.7 The arrangement of actuators", "texts": [ " The silicone currently used in this trial manufacturing is sufficient for recreating the texture of the skin. However, it loses flexibility after 1 or 2 years, and its elasticity is insufficient for large joint movements. Very humanlike movement is another important factor for developing androids. In order to realize humanlike movements, the author developed an adult android; the child android was too small to install a sufficient number of actuators in it. Figure 9.6 shows the developed adult android. The android has 42 air actuators for the upper torso except fingers. Figure 9.7 shows the arrangement of the major actuators. The android consists of rotary pneumatic actuators and linear pneumatic actuators: Rotary pneumatic actuator Hi-Rotor, Kuroda Pneumatics Ltd. (http://kuroda-precision.co.jp/KPL/english/index.html) 9 Android Science 201 Linear pneumatic actuator Air Cylinder, SMC Corporation (http://www.smcusa.com/) The specifications of the actuators are as follows. The output power can be decided by the diameter of the air cylinder and the air pressure provided by the air compressor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002886_jae-209506-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002886_jae-209506-Figure17-1.png", "caption": "Fig. 17. Distribution of magnitude of magnetic flux density for load state for 17 and 35 harmonics.", "texts": [ " These are harmonics produced by the rotor slots (32 slots from here 16 harmonics). As you can see their impact extends over a large area of the stator, hence despite the small amplitude and frequency, the losses caused by them are significant. A potential method of limiting these losses would be the use of closed rotor slots, however, this reduces the torque developed by the motor due to the reduction of the main flux. Because the main flux is still smaller compared to the low frequencies, it is impossible to get the torque we need. Figure\u00a017 however shows harmonics produced by the stator slots (36 slots). As can be seen, they are formed mainly in the tips of the rotor teeth. The increase in frequency associated with the increase of the un cor rec ted pro of ver sio n Fig. 15. Distribution of magnitude of magnetic flux density for load state for 15 and 16 harmonics. Fig. 16. Distribution of magnitude of magnetic flux density for load state for 29 and 30 harmonics. harmonic order means that a relatively small area of influence can generate noticeable losses" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002627_0954406220945728-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002627_0954406220945728-Figure4-1.png", "caption": "Figure 4. The schematic diagram of AFP machine.", "texts": [ " According to Hooke\u2019s law, the elongation of the elastic beam can be obtained L \u00bc F EA L \u00f04\u00de and from its deflection curve, the rotation angle and the maximum deflection can be obtained \u00bc M L\u00fe L\u00f0 \u00de EI \u00f05\u00de W \u00bc M L\u00fe L\u00f0 \u00de 2 2EI \u00f06\u00de thus, the transformation matrix between oxy and o0x0y0 is T F,M\u00f0 \u00de \u00bc TX W\u00f0 \u00deTY L\u00f0 \u00deRC \u00f0 \u00de \u00bc c s 0 W s c 0 L 0 0 1 0 0 0 0 1 2 6664 3 7775 \u00f07\u00de therefore T can be regarded as the pose transformation matrix generated by a joint interface. By analyzing the six motion axes of the AFP machine, it can be seen that except for X- and Caxes, the motion of the remaining four axes will cause gravity deformation. Therefore, based on the model proposed above, corresponding joint interface is replaced by an elastic beam, and the schematic diagram is shown in Figure 4, where the red parts represent the abstract of elastic beam deformation models. Analysis of the joint interface between A- and B-axis According to kinematic chain, the deformation of the joint interface between A- and B-axis is affected only by the motion of A-axis. Thus it can be simplified to a twodimensional situation, as shown in Figure 5(a), where the blue line represents elastic beam deformation model and its length is LB, the elastic modulus is EB, the crosssectional area is AB, and the moment of inertia is IB, the circle H represents the fiber placement head of weight Gh" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002854_j.engfailanal.2020.105028-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002854_j.engfailanal.2020.105028-Figure9-1.png", "caption": "Fig. 9. Safety factor of connecting rod.", "texts": [ " But the piston of the engine does reciprocating movement for a long time, and the connecting rod is affected by alternating stress, resulting in high cycle fatigue. According to the maximum running speed of the engine 2700 rpm, the certification test time is 150 h. Because it is a two-stroke engine, the connecting rod is impacted once on every revolution of crankshaft. So, the number of alternating impacts on the connecting rod is 2.43x108. In ANSYS software, Goodman equation is set to solve high cycle fatigue, and the fatigue safety factor is shown in Fig. 9. It can be seen from Fig. 9 that the fatigue safety factor in the red area of the connecting rod is relatively small. Mark the red area as point 1 and 2, the safety factor of point 1 is 0.9 and the safety factor of point 2 is 0.88. Both safety factors are less than 1, which is easy to lead to fatigue in the engine life cycle. Using Goodman method to solve the fatigue safety factor of connecting rod, as shown in Fig. 10. \u03b4fs is the maximum alternating stress when the mean stress is 0. Among this, \u03b4mis the mean stress, \u03b4y is the yield strength of the part material and \u03b4TSis the tensile strength of the material" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002569_s00170-020-05485-6-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002569_s00170-020-05485-6-Figure9-1.png", "caption": "Fig. 9 Experimental setup inside of LASERTEC 65 3D", "texts": [ " Three hundred sixteen-liter stainless steel was chosen for the powder feedstock and substrate material for its popular use in manufacturing and relatively low cost. The metal powder used in this study is MetcoClad 316 L-Si manufactured by Oerlikon Metco [22]. Its chemical composition is shown in Table 5. Its particle size distribution is shown in Table 6. The substrates were cut from hot rolled, annealed bar stock to 74 \u00d7 74 \u00d7 6 mm, and then ground to ensure that they were completely flat and level. The experimental setup inside of the machine is shown in Fig. 9. An in-house induction heating system was used for the experiments. All components were purchased off-theshelf except for the induction coil and electronics housing, which were fabricated [23]. The heater system consists of a DC power supply, a zero-voltage switching (ZVS) inverter, a pancake induction coil made of hollow copper tubing, a water-cooling system for the coil, and a vortex tube spot cooler for the inverter electronics. An induction coil was placed directly beneath the substrate for preheating" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.9-1.png", "caption": "Fig. 3.9 In describing a combined horizontal\u2013vertical movement, one has to distinguish clearly if the vertical movement is (a) a rotation about the space-fixed y-axis sy, which remains fixed, or (b) a rotation about the object-fixed y-axis by, which moves with the object", "texts": [ " For rotations about a single axis, nodistinctionhas to bemadebetween rotations about body-fixed or space-fixed axes. Since the body-fixed and space-fixed coordinate systems coincide when the object is in the reference position, the axis about which the object rotates is the same in the body-fixed and space-fixed system. But this is no longer the case for combined rotations about different axes. For such rotations the elements ofR are no longer determined by the relatively simple formulas in Eqs. (3.14)\u2013(3.16). The example in Fig. 3.9 may help to better understand the problem: how should we distinguish between a downward movement of the object by a rotation about the space-fixed axis sy (as shown in Fig. 3.9a) and a downward movement by a rotation about the rotated, body-fixed axis by (Fig. 3.9b)? 3Appendix A.3.3 contains the proof that the body-fixed representation of rotations uses the inverse (i.e., the transpose) rotation matrix compared to the space-fixed representation. 3.4 Combined Rotations 39 Mathematically, the difference between rotations in space-fixed coordinates and body-fixed coordinates lies in the sequence in which the rotations are executed. This is illustrated in Fig. 3.10. The upper column (Fig. 3.10a) shows a rotation of an object about sz by \u03b8 = 90\u25e6, followed by a rotation about the space-fixed axis sy by \u03c6 = 90\u25e6", "10a and c demonstrate that rotations about space-fixed axes and rotations about object-fixed axes in the reverse sequence lead to the same final orientation. And Eq. (3.22) is equivalent to a rotation about bz by \u03b8 , followed by a rotation about the body-fixed axis by by \u03c6. A mathematical analysis of this problem can be found in (Altmann 1986). This can be summarized as Rule 2: A switch from a representation of subsequent rotations from space-fixed axes to body-fixed axes has to be accompanied by an inversion of the sequence of the rotation matrices. This also gives the answer to the problem raised byFig. 3.9: the combination of two rotations about the space-fixed axes sz and sy, as shown in Fig. 3.9a, ismathematically described byEq. (3.20), while the combination of two rotations about the object-fixed axes bz and by, as shown in Fig. 3.9b, is described by Eq. (3.22). Rotations about space-fixed axes are often called \u201crotations of the object\u201d or \u201cactive rotations\u201d, since in successive rotations only the object is rotated, and the axes of the successive rotations are unaffected by the preceding rotations of the object. Rotations about object-fixed axes are often referred to as \u201crotations of the coordinate system\u201d or \u201cpassive rotations\u201d, since each rotation changes the coordinate axes about which the next rotations will be performed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002952_icee50131.2020.9261035-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002952_icee50131.2020.9261035-Figure5-1.png", "caption": "Fig. 5. The schematic of the rotary inverted pendulum.", "texts": [ " The plant is consisted of two computers, a DAC card and the rotary inverted pendulum. The host computer is connected to the target computer through an Ethernet cable. The controller is built in the host computer and implemented to the system through the target computer and DAQ card. Authorized licensed use limited to: Tsinghua University. Downloaded on December 19,2020 at 11:32:14 UTC from IEEE Xplore. Restrictions apply. Fig. 2. An overview of system installation. Fig. 3. Motor driver and the DAQ card Figure. 5 presents the rotary pendulum schematic diagram: Where \u03b8 is the angle of arm and\u03b1 angle of the pendulum deviating from the upright position with radian unit. , , ,\u03b8 \u03b8 \u03b1 \u03b1 are the arm and pendulum angular velocity and angular acceleration respectively. The physical definition and unit of parameters are listed in Table 1. TABLE I. THE PHYSICAL DEFINITION OF PARAMETERS OF THE SYSTEM Parameter Definition Parameter Definition m Integrate mass of the pendulum g Gravity coefficient M Integrate mass of the arm T Control input torque 1m Mass of the arm 1I Moment of Inertia of the arm 2m Mass of the pendulum 2I Moment of Inertia of the pendulum 1l Length of arm B\u03b1 Viscous damping constant of pendulum Authorized licensed use limited to: Tsinghua University" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000381_icra.2019.8793259-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000381_icra.2019.8793259-Figure2-1.png", "caption": "Fig. 2: Robot body, world, and estimated world reference frames.", "texts": [ " Touchdown (TD) is the transition between flight and stance when the foot strikes the ground. Liftoff (LO) is the transition between stance and flight when the foot leaves the ground. In flight, a jumping robot has little control over its center of gravity (CG) trajectory without specialized means to apply large forces in the air. Neglecting drag, its CG trajectory in flight follows a ballistic parabola. The robot estimates its orientation relative to a world-fixed reference frame W with basis wx, wy , wz as shown in Fig. 2. The robot can equivalently be considered to be estimating the world frame\u2019s orientation relative to the robot\u2019s frame B. We denote vectors in W\u0302 , with a hat. The following sections detail our SLIP Hopping Orientation and Velocity Estimator (SHOVE). This estimation scheme attempts to correct roll and pitch attitude errors immediately following LO. When the robot detects LO, it makes the following computations: 1) Calculate liftoff velocity using model of leg dynamics 2) Compute liftoff velocity error by comparing expected velocity from controlled touchdown attitude with 1) 3) Compute attitude corrector values for roll and pitch using liftoff velocity error 4) Subtract correction values from attitude estimate This section describes our attitude corrector algorithm for the roll and pitch of SLIP-like hopping robots" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001620_s12206-015-0908-1-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001620_s12206-015-0908-1-Figure7-1.png", "caption": "Fig. 7. Static stress and strain state under cruise cycle.", "texts": [], "surrounding_texts": [ "During the operation of an aero-engine, the turbine disc is subjected to a mixed load: centrifugal forces, thermal stresses, aerodynamic forces and vibratory stresses. Of course, high speed results in large centrifugal forces and high thermal gradients result in thermal stresses. Among them, the aerodynamic forces and vibratory stresses have little effect on the static strength of the turbine disc. Therefore, when analyzing the turbine disc with finite element method, the centrifugal forces and thermal stresses are the main consideration. The speed spectrum of the turbine disc is determined by the flight mission, and it consists of three parts [11]: low frequency cycle, full throttle cycle and cruise cycle. Any speed spectrum can be considered as a combination of these three basic cycles. The speed spectrum of the turbine disc is shown in Table 3. The temperature spectrum is derived based on the measurement data. In this study, the temperature spectrum of the turbine disc was loaded on the three-dimensional model by ANSYS parametric design language. For each basic cycle mentioned above, there are 100000 temperature data points of the turbine disc. Table 4 shows part of the temperature data points under full throttle cycle, where X, Y and Z represent the coordinate value of a point of the three-dimensional model." ] }, { "image_filename": "designv11_22_0001243_s00170-015-6914-8-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001243_s00170-015-6914-8-Figure4-1.png", "caption": "Fig. 4 Model presented for the motion of cylindrical nanoparticle in the spinning mode", "texts": [ " In order to model the spinning mode of motion, the contact forces, which are present in the system as friction, are taken into consideration; this includes deformations in particles, and by minimizing the energy of the system, the rotation axis can be specified based on the location where the force is applied. To determine the friction energy, the nanotube is divided into two sections. Point X0 denotes the location of the rotation axis and X1 is the location where the pushing force is applied, which has been considered as a variable, according to Fig. 4 (for demonstration purpose, cantilever\u2019s tip is considered as a sphere). Now, if we want to obtain the work done by the friction force at the two sides of the rotation place, we will have [3, 4]: T ab le 1 G eo m et ri ca ls pe ci fi ca tio ns an d m ec ha ni ca lp ro pe rt ie s of ca nt ile ve r Pr oc es s co ns ta nt s M ec ha ni ca lp ro pe rt ie s G eo m et ri ca ls pe ci fi ca tio ns D is ta nc e of pr ob e tip al on g th e z (Z L ) C on ta ct lo ca tio n (X ) C on ta ct an gl e (\u03c6 ) D en si ty (\u03c1 ) P oi ss on \u2019s ra tio n (\u03bd ) E la st ic m od ul us (E ) P ro be he ig ht (H ) C an til ev er th ic kn es s (t ) C an til ev er w id th (W ) C an til ev er le ng th (L ) Z L \u00bc H \u00fe L si n :7\u00f0 \u00de Z L \u00bc 1: 47 10 \u22125 0", " Then, following the integration, we will get: W rotate \u00bc 1 2 f \u0394y x1\u2212x0\u00f0 \u00de L2\u22122Lx0 \u00fe 2x0 2 \u00f020\u00de In order to obtain x0, since x1 is constant, we take its derivative with respect to x0. After differentiating and setting the resulting relation equal to zero, and considering x 0 1 \u00bc x1 L , we will have [3, 4]: x 0 0 \u00bc x 0 1 ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x 0 1 2\u2212 1 2 2x 0 1\u22121 r \u00f021\u00de Using the above equation, spinning point is introduced. Considering geometrical relations (Fig. 4), it is obvious that after a certain point particle and probe are detached so based on geometrical relations a loop is introduced to find time of separation. Now by using the presented relations and the geometry of the nanoparticle (Fig. 4), the ultimate angle of rotation is specified. In below equations, \u03b8ult is ultimate angle of rotation, Ti is the second critical time, Vsub is substrate speed, L is particle\u2019s length, x1 is the distance of pushing point, S1, S2,and D are geometrical parameters which are shown in Fig. 4. \u03b8ult \u00bc Ti V sub L\u2212 x 0 1 \u00fe ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi x 0 1\u2212 1 2 2x 0 1\u22121 r ! \u00f022\u00de S2 \u00bc D\u2212Dcos \u03b8ult\u00f0 \u00de \u00f023\u00de S1 \u00bc L\u2212x1 \u00f024\u00de Using Eqs. 22 to 24 in a loop, the ultimate rotation angle and second critical time are determined. In the first step, an initial value of 0.01 is given to Ti so with respect to a constant pushing point (x1) and defined particle length (L), we have \u03b8ult. In the next step, S2 which is a geometrical parameter and is shown in Fig. 4 is defined. In the final step, S1 which is a constant value based on relation 24 is compared with S2. While S1 is bigger than S2, an interval value of 0.1 is added to Ti and Eqs. 22 to 23 are reevaluated in a loop. When S2 is bigger or equal to S1, the ultimate rotation angle and second critical time are determined. For a 1.8-\u03bcm-long particle with a radius of 60 nm, a final rotation angle of 67\u00b0 is obtained within 11.3 s. These findings agree with the presented experimental results [6]. At this point, by knowing the dynamic motion modes, the manner of movement of cylindrical particles is explored, and the critical force and time, which are the main parameters of the manipulation operations, are obtained" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000754_1.4031440-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000754_1.4031440-Figure1-1.png", "caption": "Fig. 1 Gas bearings for oil-free TCs: (a) bump-type foil bearing [12] and (b) FPTPGB [13]", "texts": [ " GFBs offer distinctive advantages over rolling elements bearings and rigid surface gas bearings including reliable high-temperature operation with enhanced damping and large Contributed by the Structures and Dynamics Committee of ASME for publication in the JOURNAL OF ENGINEERING FOR GAS TURBINES AND POWER. Manuscript received July 13, 2015; final manuscript received August 16, 2015; published online October 13, 2015. Editor: David Wisler. Journal of Engineering for Gas Turbines and Power APRIL 2016, Vol. 138 / 042501-1 Copyright VC 2016 by ASME Downloaded From: http://gasturbinespower.asmedigitalcollection.asme.org/ on 01/20/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use tolerance to debris and rotor motions. Figure 1(a), taken from [12], depicts a typical corrugated bump-type foil bearing which is a compliant surface hydrodynamic bearing using ambient air as the working fluid media. The bump strip layers are an elastic support with engineered stiffness and damping characteristics. Note that the static load capacity and dynamic forced performance of GFBs depend largely on the material properties and geometry of the support elastic structure. In addition, they also depend not only on the surface conditions and load applied but also on the environmental conditions (i.e., ambient pressure and temperature, atmospheric chemical composition and humidity, etc.). Tilting pad bearings are commonly used in high-performance turbomachinery because of their superior stability characteristics due to negligible cross-coupled stiffnesses. Figure 1(b), taken from [13], shows a FPTPGBs which are machined as a single piece using wire electrical discharge machining process. The bearing offers a compact unit which is easier to install and maintain while providing similar rotordynamic advantages as a tilting pad bearing. FPTPGBs also eliminate pivot wear, contact stresses, and pad flutter, and minimize the manufacturing tolerance stack-up. If machined with a radial stiffness, this bearing type gives a margin of tolerance for shaft growth and permits shaft excursions beyond its nominal clearance" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001428_1350650114559617-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001428_1350650114559617-Figure1-1.png", "caption": "Figure 1. Bearing load distribution (dotted line indicates the state before stress).", "texts": [ " In this paper, the cylindrical roller bearing was taken as the research object where the mechanical analysis a deformation coordination model has been established, which can calculate the rollers\u2019 stress when geometry errors exist. Each roller\u2019s force has been calculated by the numerical iteration method and the influences of several different size errors to the rolling bearing stress have been studied and analyzed. Analysis of Roller force in rolling bearing A cylindrical roller bearing is taken as the research object. As shown in Figure 1, without considering roller geometrical error, the rollers in the lower half ring are under force whereas the upper half are not when bearing is under radial load. According to the derivation by Stribeck, the stress of each roller can be calculated by the deformation coordination and force balance which is given as Fr \u00bc Q0 \u00fe 2Q1 cos\u2019\u00fe \u00fe 2Qk cos 2\u2019 \u00f01\u00de As showing in Figure 1 and according to deformation coordination, the following equations are obtained 0 \u00bc r \u00f02\u00de According to the Palmgren\u2019s derivation k / Q0:9 k \u00f04\u00de Therefore, the following equation is obtained Fr \u00bc Q0 1\u00fe 2cos1:9\u2019\u00fe \u00fe 2cos1:9k\u2019 \u00f05\u00de In Figure 1, Q0 and all the other rollers\u2019 loads can be calculated given the Fr. But this method is just a rough calculation according to the elastic deformation coordination without considering roller diameter error, and it has a great limitation where it can only be applied in the roller symmetry distribution, and also it cannot calculate the deformation of each roller. Calculation method with considering roller geometry error Elastic contact analysis. When two cylindrical surfaces are in contact with each other by force, ideally the contact area will change from a line segment into a rectangle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002673_j.matpr.2020.07.229-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002673_j.matpr.2020.07.229-Figure3-1.png", "caption": "Fig. 3. Contour plot of 1st Mode shape (", "texts": [], "surrounding_texts": [ "While conducting the model analysis of the composite cantilever beam, convergence criteria was achieved after several iterations. Using an iterative method, we increase the number of elements by decreasing the size of elements. To determine the suitable mesh size which produces a precise result, the mesh size was decreased until the natural frequencies for a mode shape converged. For the modal analysis of the beam, the number of nodes obtained were 11,028 and the number of elements were 10,983 for an element size of 3 mm. a meshed model of the beam is shown in Fig. 2." ] }, { "image_filename": "designv11_22_0002086_tjj-2016-0066-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002086_tjj-2016-0066-Figure1-1.png", "caption": "Figure 1: Schematic of a variable speed rotor-seal model.", "texts": [], "surrounding_texts": [ "Keywords: variable speed rotor-seal, straight-through labyrinth seal, rotor vibration, leakage, aerodynamic force\nPACS\u00ae (2010). 47.85.Gj\nRotor-seal stability has always been a significant issue of modern gas turbine. Since the early 60s, Childs et al. summarized a number of studies on this interaction, showing a variety of effects on the rotor with different types of seals [1\u20134]. Experimental investigations on rotordynamic coefficients and leakage performance were conducted on the Annular Gas Seal Test Stand (AGSTS) [5, 6], and results were compared with the predictions based on a 1-CVM model derived by Childs and Scharrer [2]. The investigations [7\u20139] revealed that the stiffening or softening effect on rotor directly depended on the directions of the aerodynamic forces induced by sealing flow field. In testing,\nseal forces were indirectly measured by force sensors on the excited stringers with constant speed condition. Seal forces cannot be dynamically and directly measured by seal test rig, and rotordynamic coefficients are not available to estimate the variable speed rotor-seal stability.\nNumerical approaches were widely used in leakage and stability predictions of annular gas seals. A numerical code was compiled by Tam [10] on the basis of an altered form of the three-dimensional Navier-Stokes equation, the radial and tangential components of the seal forces can be calculated in this code, and the predicted results were in accordance with experimental results in literature [2]. Wilbur [11] developed a computational fluid dynamic (CFD) code to produce three-dimensional flow information, and this code has been used in NASA\u2019s advanced engine programs to analyze, design and optimize advanced seals. The tip clearance height and the pressure distribution of seal flow field could be obtained by this code. Self compiled programs were progressively replaced by the commercial software in numerical analysis. A full 3D, eccentric model was defined and solved by Moore [12] to improve on seal force prediction. Three-dimensional fluid dynamics methods were carried out to extract seal aerodynamic forces in literature [13, 14]. Zhang et al. [15] compared the aerodynamic forces in labyrinth and honeycomb seals, results showed that the honeycomb seal can produce significant aerodynamic force, which is in accordance with the experimental results of Childs et al. [4].\nFor most numerical investigations, leakage and seal force dependency on rotating speed was conducted with constant speed boundary condition. For example, Zhang et al. [15] incorporated the seal flow field deformation caused by rotor vibration and achieved a relative real-time interaction force result under constant speed condition, whereas these numerical approach are not available to variable speed conditions. This limitation prevented seal forces measurement under variable speed condition, which is critical to the rotor stability of aviation gas turbine with variable speed behavior. For example, in a flight mission of three hours, variable speed behavior to occur 10%of the time.While stability of the rotor is important and critical even for short durations during the flight mission, increased leakage that deviates from ideal cases (i. e. constant speed rotor) is acceptable for short durations during the flight. *Corresponding author: Hai Zhang, Institute of Turbomachinery, College of Power and Energy Engineering, Harbin Engineering University, Harbin 150001, China, E-mail: zhanghai83821@163.com Xingyun Jia, Qun Zheng, Shuangming Fan, Zhitao Tian, Institute of Turbomachinery, College of Power and Energy Engineering, Harbin Engineering University, Harbin 150001, China\nBrought to you by | New York University Bobst Library Technical Services Authenticated\nDownload Date | 12/8/16 3:15 AM", "Thus, this study is (a) more important to complete stability concerns for variable speed behavior, and (b) important for leakage only if the variable speed behavior occurs formajority of the flight mission.\nA variable speed rotor-seal model is composed of annular seal and rotor with variable rotating speed. Type of seal is defined as the straight-through tooth-on-stator labyrinth (LABY) seal, nine straight teeth are uniformly arranged on stator, producing eight seal cavities. Figures 1 and 2 show the schematic and dimensions of labyrinth seal. The height of the initial radial clearance without rotor vibration is h of 0.2mm, the throttled tooth thickness b is constant to 0.2mm. For this variable speed rotor-seal model, rotating part of rotor-seal system in aviation gas turbines was\nsimplified to a cylindrical shaft. Rotor diameter is 116mm, and seal diameter is 126mm. The numerical values of material properties of the rotor: the density of rotor material is 7800kg/m3, the elastic modulus is 2.1 \u00d7 1011, Poisson\u2019s ratio is 0.3. Bearing stiffness is 1.75 \u00d7 107 N/m, and the damper of the bearing is regarded as 0. Rotating speed increases from 18 to 30 krpm.\nRotor undergoing vibration by precessing about its central axis at a frequency of \u03a9 and whirling around its own axis at a frequency of \u03c9. The shape of seal flow field transformed when rotor deviates its axis. The tooth tip clearance region becomes narrow and slight when rotor precesses to seal, wide and obvious for the contrary side. Thus, the rotor precession effect on seal must be incorporated in the fluid dynamic calculations.\nEach elliptical whirling trajectory of rotor can be divided into a positive and a negative precession trajectory. Positive precession force is an unbalanced mass force acting on the positive precession trajectory without the negative one. For constant rotating speed condition in literature [15], vibration rotor follows a circle precession trajectory, and flow field area contains constant. However, both locations and area of seal flow field are turned into non-uniform and time-dependent in a variable speed rotor-seal model, therefore the seal inner flow situation becomes not an absolute cyclical change but a dynamic procession. The transient trajectory of variable rotating speed rotor was calculated by dynamics analysis in ANSYS and self compiled code, and then dynamic vibration characteristics (vibration frequency and amplitude) were incorporated in the fluid dynamic calculation as transient boundary conditions. Process diagram for the solution is set out in Figure 3.\n2 X. Jia et al.: Variable Speed Rotor-Labyrinth Seal System\nBrought to you by | New York University Bobst Library Technical Services Authenticated\nDownload Date | 12/8/16 3:15 AM", "Rotor natural frequency and vibration frequencies were obtained from Campbell chart [16]. Critical angular velocity is the rotating angular velocity, which is equal to the precession angular velocity. Thus, the critical angular velocity and critical speed of the rotor can be determined according to the equation \u03a9=\u03c9 (where \u03a9 is vibration frequency and \u03c9 is rotational speed) and rotor precession angular velocity and frequency.\nFor rotordynamic code, the gyroscopic effect [G], and the damping effect [B] were incorporated in the general dynamic equation leading:\nM\u00bd \u20acU + G\u00bd + C\u00bd \u00f0 \u00de _U + B\u00bd + K\u00bd \u00f0 \u00de Uf g= ff g (1)\nwhere [M], [C] and [K] are mass, damping and stiffness matrices, respectively, and {f} is an external force.\nBoth the strong vortices in sealing cavities and jet flow in teeth tip clearances are representative threedimensional flow so that a compressible ReynoldsAveraged Navier-Stokes is applied in the computational fluid dynamics calculation. A commercial multi-purpose CFD code, ANSYS CFX 14.5 software, is used to predict the flow situation in seal. Transient mode is utilized in CFX due to the sealing flow requires real time information to describe flow characteristics. Note that for variable speed rotor, the time step for transient calculation is set to meet the requirements of eq. (2):\nH Vmax >Timestep (2)\nWhere H is the height of the first layer mesh and Vmax is the maximum value of moving velocity of the first layer mesh. Working fluid is modeled as ideal gas and turbulence model k \u2212\u03c9 is applied to simulate the turbulence characteristic. Method of numerical and boundary conditions are described in Table 1.\nMonitoring points P1-P8 are defined averagely in each circle of precession trajectory, as shown in Figure 4. The\nsequence of these points follows the rotor precession. The results for these points were assumed to represent the dynamic characteristic through the whole cycle. The clearance between seal tips and stator were exaggerated to observe variations in the small clearance more easily.\nThe numerical predicted straight-through labyrinth seal performance and the published CFD results of straight-through labyrinth seal tested by Yahya, et al. [17], and results obtained by the most commonly referenced 1D labyrinth seal leakage correlation (Egli\u2019s correlation) in literature [18, 19] were compared to validate the accuracy and reliability of numerical methods and transient CFD variable speed rotor-seal model. Leakage with respect rotor speed from 0 to 80 krpm were plotted in Figure 5.\nThe shape and critical dimensions of labyrinth rotorseal model, boundary conditions for numerical method were consistent with the test seal and operational conditions. Leakage across variable speed conditions from 10 to 40 krpm were chosen to validate the numerical method and transient CFD rotor-seal model under 0.4 Mpa pressure difference. It is seen that the leakage calculated by using the present CFD procedure is compatible with the leakage derived from testing and Egli\u2019s correlation.\nA computational whole circumferential field and mesh can capture real-time information of variable speed\nTable 1: Method of numerical and boundary conditions.\nInlet total pressure . Mpa, . Mpa and . Mpa Inlet temperature (\u00b0C) Outlet static pressure . MPa Rotational speed \u2013 (krpm) Discretization scheme High resolution Computational method Time marching method Mesh deformation Dynamic mesh, increase quality near boundary Turbulence model k \u2212\u03c9 Fluid Air (ideal gas) Wall properties Adiabatic, smooth surface\nX. Jia et al.: Variable Speed Rotor-Labyrinth Seal System 3\nBrought to you by | New York University Bobst Library Technical Services Authenticated\nDownload Date | 12/8/16 3:15 AM" ] }, { "image_filename": "designv11_22_0000043_s12541-019-00047-7-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000043_s12541-019-00047-7-Figure13-1.png", "caption": "Fig. 13 The dynamic model of the gear system", "texts": [ " One case is only that the excitation of time-varying mesh stiffness exists; the other is where two excitations are both considered. The results of two cases are expressed in time and frequency domains, phase plan plots (14)fs(t) = Fs sin ( st ) ( 0 \u2264 t \u2264 ts ) (15)\u222b ts 0 fs(t)dt = mevs (16)ts = 2 mevs Fs 1 3 and Poincar\u00e9 maps. Furthermore, we use the wavelet packet analysis to extract the impact feature. Through the comparison of two cases, we investigate the effect of the meshing-in impact on the dynamic behaviour of gear system, considering SRNPE. Figure\u00a013 depicts a dynamic model of a helical gear pair. In this picture, the pinion is left-handed and rotates about the z-axis counter clockwise. Both gears are treated as rigid discs, which are connected to each other by a time-varying mesh spring km in the plane of action along the normal direction. 6 DOFs for the pinion and the gear are included in the dynamic model. They are translation motion along the x-, y- and z-axes, swinging motion around the x-, y-axes and torsional motion around the z-axis", "\u00a0(19), \u03b4 is the relative displacement between the gear pair along the line of action and is expressed as follows: where \u03c8 is the angle between the plane of action and the positive y-axis, which is defined as: (18)cm = 2\ud835\udf09 \u221a\u221a\u221a\u221a k\u0304mIzpIzg Izpr 2 bg + Izgr 2 bp (19) = [( xp \u2212 xg ) \u22c5 sin + ( yp \u2212 yg ) \u22c5 cos + uzp + uzg ] \u22c5 cos b + [( uxp + uxg ) \u22c5 sin + ( uyp + uyg ) \u22c5 cos \u2212 zp + zg ] \u22c5 sin b, (20) = { t \u2212 , counterclockwise rotation for pinion t + , clockwise rotation for pinion , 1 3 where \u03b1t is the operating pressure angle, and \u03c6 is the position angle of the centre line in Fig.\u00a013. Case 1 Only the excitation of time-varying mesh stiffness. In this case, we use only TCA and LTCA to compute the loaded transmission error without SRNPE in a meshing period. Then, we convert the loaded transmission error to the tooth elastic deformation by Eq.\u00a0(5). The ratio of the normal force and the tooth deformation is the time-varying mesh stiffness, which is determined using Eq.\u00a0(21). where km is the time-varying mesh stiffness; \u0394x is the tooth deformation; Fn is the normal force; Tg is the load torque, (21) { km = Fn \u0394x Fn = Tg rbg \u22c5 cos b , with Tg= 500\u00a0N\u00a0m; rbg is the base radius of the gear; and \u03b2b is the base helix angle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002798_s00170-020-06152-6-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002798_s00170-020-06152-6-Figure11-1.png", "caption": "Fig. 11 Combination of rectangular datum target areas (B1,2) and inspection oriented datum target feature (C1)", "texts": [], "surrounding_texts": [ "The well-known \u201cmaximum utilization\u201d rule concerning datum hierarchy states that \u201cIf a Datum Feature can and may constrain a degree of freedom, it must.\u201dWhen a DF is used to establish a datum, it constrains, or locks, some degrees of freedom of an ideal feature (e.g., the tolerance zone). As per ISO 5459, the maximum number of degrees of freedom that can be constrained by this integral feature is equal to, or less than, six minus the invariance degree of the nominal integral DF. However, high level of flexibility is again here provided to the designer, concerning the designation of the number of locked or released degrees of freedom in each datum of a datum system. This can be specified by the assignment of complementary indication ([PL], [SL], [PT], ><, [ ] or [Tx,Ty,Tz,Rx,Ry,Rz]) added after the datum identifier symbol in the relevant datum section, as shown in the example given in Fig. 7." ] }, { "image_filename": "designv11_22_0000133_b978-0-08-102814-8.00013-5-Figure11.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000133_b978-0-08-102814-8.00013-5-Figure11.1-1.png", "caption": "FIGURE 11.1", "texts": [ " have been synthesized. Sol gel synthesis, hydrothermal or solvothermal growth, physical or chemical vapor deposition, low-temperature aqueous growth, chemical bath deposition, or electrochemical depositions are methods used for nanoscale material formation. Electronic materials include insulators, semiconductors, conductors, and superconductors. A solid insulator is a substance with a very low electrical conductivity and there is a considerable energy gap before an empty orbital becomes available (Fig. 11.1A). A semiconductor is a substance which has electrical conductivity that increases with increasing temperature, and at room temperature, the conductivities are typically intermediate between conductors and insulators (Enhessari, 2013; Khanahmadzadeh et al., 2015; Zare et al., 2009; Enhessari et al., 2010, 2012a,b, 2013, 2016a,b; Nouri et al., 2016). Semiconductors are classified as intrinsic or extrinsic semiconductors; the former band gap is very small (Fig. 11.1B), therefore, the energy of thermal motion results in jumping of some electrons from the valence band into the empty upper 358 CHAPTER 11 Application of (mixed) metal band, and the former is a semiconductor in the presence of impurities; p-type doping atoms remove electrons from the valence band and n-type doped atoms supply electrons to the conduction band (Fig. 11.1C). A conductor is a substance with an electric conductivity that decreases with increasing temperature, and has zero band gaps (Fig. 11.1D). A superconductor (classified into three types) is a class of materials with zero electrical resistance. The origin of this class of materials is related to electron phonon coupling and the resultant pairing of conduction electrons (Askeland and Phule, 2006). A metallic conductor, a semiconductor, and a superconductor can be distinguished based on the temperature dependence of the electrical conductivity. This variation can be observed in Fig. 11.2. The structure of a typical (A) insulator, (B) intrinsic semiconductor, (C) extrinsic semiconductor, (D) conductor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002801_s11661-020-06013-7-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002801_s11661-020-06013-7-Figure2-1.png", "caption": "Fig. 2\u2014The side and front views of the three-layered specimen with a polished surface for in-situ observation during uniaxial compression test.", "texts": [ ", Taiwan). The compression rate was 0.125 mm/min. The experimental procedures and the compression rate followed the ASTM E9-19 standard. For both the SLM and HIP conditions, three specimens were tested and the results are presented. To observe the fracture evolution during the compression test, a polished surface of a specimen was examined under a CCD image recording system with six million pixels for in-situ observation. The details of the specimen for in-situ compression test are illustrated in Figure 2. The images of the samples were taken once every second during the in-situ test. The DIC technique, a versatile method of obtaining the strain mapping in a material after stressing, can provide valuable information about the deformation behavior. Unfortunately, the DIC technique has rarely been used in the field of AM cellular alloys to date.[8,32\u201335] In this study, two-dimensional DIC software (Vic-2D, Correlated Solutions, Inc.) was used to analyze the successive pictures obtained from the in-situ test and to calculate the strains in each of the areas of the specimens" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure17-1.png", "caption": "Fig. 17. Locations of strain gauges 5 -8.", "texts": [], "surrounding_texts": [ "The reliability of the structural analysis was evaluated ex- perimentally. The stress level at the same loading condition was compared analytically and experimentally at the same position of the test rig. The applied loading was an extreme load case for a pitch bearing. A strain gauge was used to measure the stress level of the test rig. Ten strain gauges of two different types were applied at key locations in the test rig, as shown in Figs. 16-18. The gauge type was chosen according to the shape of the mounted parts, and the specifications of each type are shown in Table 3. Fig. 19 shows the stress results of the different approaches. The analytical and experimental results show good agreement in terms of the tendency and magnitude of stress. Therefore, it may be concluded that the analysis method is suitable for estimating the test rig\u2019s stress level." ] }, { "image_filename": "designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure5-1.png", "caption": "Fig. 5. SARRUS 3H3h mechanism.", "texts": [ " (26) beginning with the 3rd-order, k\u0302 T j D A\u2032\u2032k = 2k\u0302 T j D A\u2032\ufe38 \ufe37\ufe37 \ufe38 0 A+A\u2032k, k\u0302 T j D A\u2032\u2032\u2032k = 3k\u0302 T j D A\u2032\u2032\ufe38 \ufe37\ufe37 \ufe38 0 A+A\u2032k + 3k\u0302 T j D A\u2032\ufe38 \ufe37\ufe37 \ufe38 0 A+A\u2032\u2032k \u2212 6k\u0302 T j D A\u2032\ufe38 \ufe37\ufe37 \ufe38 0 A+A\u2032A+A\u2032k, ... (52) it is seen that the right hand sides of these equations vanish. Thus the vanishing products 0 = k\u0302 T j D A\u2032 are also sufficient but not necessary conditions for finite mobility, see also SUGIMOTO [26] and MILENKOVIC [13]. Example: 3H3h mechanism (SARRUS). An example for the special solution (51) is the SARRUS linkage with six helical joints. This special 6H mechanism has two groups of three adjacent parallel joint axes, according to Fig. 5. Five screw axes are given by A = [ a1 a1 a1 a4 a4 ae1 ae2 ae3 ae4 ae5 ] (53) and the sixth screw axis by a\u0302T 6 = [ a4 T ae6 T ] . The reciprocal screw obtained from the reciprocity condition in Eq. (25) is k\u0302 = [ 0 a\u03031 a4 ] . (54) To proof the fulfilment of the 2nd-order mobility condition from Eq. (26), k\u0302 T D A\u2032 k = 0, (55) matrix A\u2032 = [ 0 a\u0302\u2032 2 . . . a\u0302\u2032 5 ] is calculated by means of Eq. (18). The dual vector product of the parallel screws a\u03021, a\u03022, a\u03023 with a1 = a2 = a3 is given by \u02dc\u0302a1 a\u0302i \u2261 [\u0303 a1 ae1 ] [ a1 aei ] = [ 0 b1i ] , i = 2, 3, (56) with the common normal vector b1i from a\u03021 to a\u0302i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002569_s00170-020-05485-6-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002569_s00170-020-05485-6-Figure2-1.png", "caption": "Fig. 2 Simulated substrate and clad placement (left) with mesh (right)", "texts": [ " The simulations modeled a single track, single layer deposition. The model consisted of a substrate, a clad, a moving surface heat flux to represent the laser, surface film conditions to model heat transfer between the materials and surroundings, and an initial temperature condition on the substrate to represent the preheating. The substrate was modeled as a solid 74 mm square by 6 mm thick part. The clad was modeled as a solid 54 mm by 3 mm by 1 mm thick section that was positioned on top of the substrate as shown in Fig. 2. The clad was divided into 18 sections, each Table 1 Material properties for 316-L stainless steel [13, 14, 16] Temperature (\u00b0C) Thermal conductivity (W/m\u2219K) Density (g/cm3) Expansion coefficient (10\u22126/K) Latent heat (kJ/kg) Specific heat (J/kg\u2219K) 2000 \u2013 6.6 \u2013 \u2013 \u2013 Initial substrate temperature (\u00b0C) 3 mm square, and was initially inactive. A linear tetrahedral mesh with 5836 nodes and 24,311 elements was used with an element length of 5 mm on the substrate and an element length of 0.5 mm on the clad section as shown in Fig. 2. The mesh size for the clad section was chosen because a convergence test was performed, and the change in the temperature results at any given node was less than 1% between a 0.5 and 0.25 mm mesh. Results were only obtained for the locally meshed clad section. Three hundred sixteen-liter stainless steel was used for both the substrate and clad materials. Material properties for 316-L stainless steel that were used for both the substrate and clad materials are shown in Table 1. The solidus and liquidus temperatures of 316-L stainless steel are 1385 \u00b0C and 1450 \u00b0C, respectively [13, 14, 16]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001851_s12206-016-0614-7-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001851_s12206-016-0614-7-Figure9-1.png", "caption": "Fig. 9. Multi-body dynamic model of the taper leaf spring of tandem suspension.", "texts": [ " The specific value of parameters in Coulomb friction could not copy the ones from other similar documents because friction is associated with many factors, such as the relative velocity, acceleration and displacement, lubrication condition, and the condition of contact surface [16, 17]. Moreover, the particularity of interleaf friction in the taper leaf spring of tandem suspension should be taken into account. Therefore, all parameters about Coulomb friction in this study were found through trial and error. Table 1 shows the specific value of parameters of Coulomb friction. The modified model was developed through the above described process. Fig. 9 shows the multi-body dynamic model of the taper leaf spring of tandem suspension. A series of tests were conducted to verify the modified model. The taper leaf spring of tandem suspension used for the tests, as shown in Fig. 10, had four leaves, 90-mm width, and non-uniform thickness. Fig. 11 is a schematic diagram of the apparatus for the tests. Fig. 11 shows that the hydraulic actuator, which is equipped with high-precision sensor in the head and controlled by computer, was used to apply the sinusoidal alternating load to the taper leaf spring of tandem suspension" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000709_psc49016.2019.9081455-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000709_psc49016.2019.9081455-Figure6-1.png", "caption": "Fig. 6. cylindrical configuration of Inductrack suspension [6]", "texts": [ " Other methods of suspension Although EDS and EMS are currently the only types of suspensions which are used in practical Maglev trains, loss of the suspension DC current can lead to sudden fall of train and therefore, catastrophic accidents [4,5]. To solve this problem, Inductrack is proposed. Inductrack topology is very similar to EDS, but as demonstrated in Fig. 5, instead of superconductive coils, Halbach array NdFeB PMs are used. Also, instead of 8- shaped coils, a track with short circuited rods is used. In this technology, similar to EDS, in lower speeds wheels are needed to support the train\u2019s weight. Inductrack suspension can be rearranged into a cylindrical configuration as shown in Fig. 6. This configuration is discussed in [6]. Another way of suspending trains is by taking advantage of the fact that when PM and superconductive materials are near each other, a repulsive force exists between them. In this design which is shown in Fig. 7, a Halbach array of PM materials are placed on the track and a set of superconductive YBCO matterial is placed on the train facing the track [7]. III. TRACTION OF MAGLEV TRAINS The best way of moving maglev trains forward is through the concept of linear electric machines" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000912_9781119016854.ch1-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000912_9781119016854.ch1-Figure7-1.png", "caption": "Figure 7. A) Model of the machining of airfoils produced by conventional subtractive processing of forged and heat treated Alloy 718 material. B) Schematic of the approach to integrate the design of material and process definition with component design optimization.", "texts": [ " As noted earlier relative to DA718, computational materials modeling has become enabling to achieve the desired microstructure and subsequent mechanical properties on a component location-specific basis. [16, 17] In addition to microstructure and mechanical property models, additional modeling and simulation capabilities have been established to support existing and emerging manufacturing processes. Machining modeling has advanced to the point where optimal machining parameters can be designed computationally to enable desired component configuration and surface characteristics to be demonstrated the first time in actual production. Figure 7 shows an example of a model for the optimization of machining of an integrally bladed rotor (IBR). In addition to conventional subtractive processes to produce final component configurations, the emerging technology of additive manufacture has been applied extensively to Alloy 718. For these processes, Alloy 718 and variants are produced into powders that are recombined into desired form through localized melting and systematic and sequential build-up. Powder-bed and powder-feed processes have been successful in producing many unique configurations, often not possible by conventional processes or that would require high cost conventional processing" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001334_amc.2014.6823326-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001334_amc.2014.6823326-Figure5-1.png", "caption": "Figure 5. An Acrobot", "texts": [ " It will be demonstrated in this paper that the convergence rate of the PSO algorithm is improved by improving its structure and making it iterative(iPSO) and it will be reasoned that the iPSO algorithm can provide convergence to the global minimum, this differs from the traditional PSO which is known to sometimes get stuck in a local minimum. Further originality can be found in the combination and implementation of this iPSO algorithm with a MPC algorithm on a GPU. Finally simulations will be used to demonstrate the proposed algorithms ability to make the Acrobot(inverted double pendulum with an underactuated base as shown in fig. 5) swing up to its upright position in real-time. In [11] the swingup task of the Acrobot was experimentally demonstrated by using a energy pumping strategy combined with a LQR controller to maintain balance at the upright posture. However the energy pumping method has limited usefulness in general problems. In recent years a push towards general algorithms that can solve complex nonlinear tasks has been made. LQRTrees has been proposed[6] and experimentally verified on an Acrobot[9], it shows an ability to execute impressively complex tasks, however a large part of the computation needs to be carried out offline in preparation for the task, this is not practical for robots that will operate autonomously in unknown environments" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002775_j.triboint.2020.106669-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002775_j.triboint.2020.106669-Figure1-1.png", "caption": "Fig. 1. The schematic diagram of the roller-ring test rig, (a) the physical system, (b) the principle of the loading system, (c) the detailed structure of the tribo-pair.", "texts": [ " A roller-ring tribo-pair is built with a real ring and profiled roller from the object bearing. Moreover, wide range conditions for the fundamental sliding-rolling phenomena are considered and not limited to a bearing. Time-Frequency analysis is conducted to explore the C. Xu et al. Tribology International 154 (2021) 106669 vibration characteristic frequency for different slip ratios. Afterwards, a simplified finite element model of the tribo-system is built to investigate the vibration characteristic under sliding-rolling contact. As shown in Fig. 1, the roller-ring tribo-system is employed for friction and loading tests. Referring to the NU209EM bearing, the \u03c612 mm roller is profiled with an additional holding section and the \u03c6108 mm ring is directly dismantled from the bearing. The roller is driven by a high-speed electric spindle with a flexible coupling to isolate the radial impact from the driver and the ring is fixed on a low-speed electric spindle, as can be seen in Fig. 1(a). Sliding-rolling motions are obtained by setting different drive speeds. The principle of the loading system and detailed structure of the tribo-pair are shown in Fig. 1(b) and (c) respectively. The ring is positioned by the supporting shaft and the roller is in contact with the ring through two supporting bearings. The loading system is mainly composed of the pillar, lever, load block, wire rope, spring, load cell, worm gear reducer and servo motor. In the test, one end of the lever is connected to the pillar and the other end is pulled by the wire rope driven by the servo motor. A bolt and a nut are fixed on the lever to apply the vertical load to the load block, supporting bearings and then to the tribo-pair" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001352_s0025654414010026-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001352_s0025654414010026-Figure2-1.png", "caption": "Fig. 2.", "texts": [ " Then \u0394A=A\u2212Aopt = \u03b12\u03c0 2\u03c9 ( m2a4\u03c94 2 + F 2 0 a2 ) \u2212 2\u03b12F 2 0 a2 \u03c0\u03c9 = \u03b12\u03c0m2a4\u03c93 4 + \u03b12F 2 0 a2 \u03c9 ( \u03c0 2 \u2212 2 \u03c0 ) . (6.7) Note that the effect of a decrease in the energy losses is realized even if the inertial properties of the mechanism are not taken into account (m = 0). This can be explained by the fact that the effective work linearly depends on the working resistance force, while the energy losses are proportional to the square of this force. As a second example, consider a conventional pumping unit used in oil industry; its general view and functional diagram are shown in Fig. 2. A specific feature of this mechanism is that its working stroke corresponds to the bar lifting with the extracted oil and no effective work is done in the free running. The following geometric characteristics of the pumping unit are chosen: the crank length is OK = 1.2 m, the connecting rod length is KL = 3.31 m, the balance arm length from the fixation point of the connecting rod to the balance arm axis is LO1 = 2.5 m, and the balance arm length from the fixation point of the rod cable to the balance arm axis is MO1 = 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000514_j.mechmachtheory.2019.103667-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000514_j.mechmachtheory.2019.103667-Figure7-1.png", "caption": "Fig. 7. Shaft dimensions and 3D representation of the face-gear drive subject of study.", "texts": [ " The amount of crowing applied to the face gear by using a shaper of 26 teeth for the gear drive with design parameters as provided in Table 1 is of about 30 \u03bcm, as shown in Fig. 6(b). The provided longitudinal crowning applied to the face gear makes the gear drive more insensitive to errors of alignment and makes this study to be aligned with actual designs of face-gear transmissions. Table 2 shows the design parameters of the pinion and face-gear shafts. Hollow shafts have been considered for both the pinion and the face gear. Both gear elements are overhung mounted outside the bearing span as shown in Fig. 7. When gear supporting shafts are not taken into account, the conditions of meshing and contact can be considered as theoretical or ideal. In this situation, the contact patterns on the gear tooth surfaces are said to correspond to the unloaded (or lightly loaded) gear drive. However, when gear supporting shafts are considered for stress analysis, real conditions of meshing and contact are simulated and the deflections of shafts under the transmitted torque will influence the contact pattern location and the level of stresses on the gear tooth surfaces" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001601_s00170-016-8487-6-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001601_s00170-016-8487-6-Figure4-1.png", "caption": "Fig. 4 a The unassembled halves of Exp-ECAE die and b the assembled die together with the tool accessories", "texts": [ " 2 h i. ffiffiffi 3 p \u00f04\u00de Experimental tests were performed using the AA6063 aluminum alloy [14]. Rods with 20 mm in diameter were purchased then machined into the billets with a diameter and length of 15 and 120 mm, respectively. The billets were initially annealed for 2.5 h at 700 K. An Exp-ECAE die with perpendicular channels (\u03d5 =90\u00b0) was used for experiments. According to Fig. 1, geometrical parameters of the die including D0, D1, r, and e were respectively measured to be 15, 23, 1, and 2 mm. Figure 4a exhibits the unassembled halves of die. The ExpECAE die together with the tool accessories are illustrated in Fig. 4b. In order to reduce the interfacial friction, MoS2 was used as lubricant. Several samples were processed through the Exp-ECAE at different temperatures and with various ram velocities while the load-displacement variations were recorded during the experiments. The microhardness of the material was measured to examine the distribution of mechanical properties and, consequently, to validate effective strain distribution determined bymeans of the FE analyses. The longitudinal section of each product, containing the symmetry plane, was first prepared and, then, the microhardness variation was measured along a specific path overlaying the deformation zone" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000843_978-3-319-18290-2_5-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000843_978-3-319-18290-2_5-Figure2-1.png", "caption": "Fig. 2. DC-DC Buck converter", "texts": [ " Note that it is possible there is another kind of singularity problem as that for the continuous-time FTSMC in [9]. That is there may exist such a case when |si(t0)| > \u03c1i and |sj(t0)| \u2264 \u03c1j for 0 \u2264 j < i \u2264 (n \u2212 1) at the initial time. Actually, in practice, to avoid this singularity problem, as suggested in [8] for the continuous-time TSMC, we can replace sj(t0) with the following mapping function: Map(sj(t0)) = { sj(t0), for |sj(t0)| > \u03c1j ; \u03b4, for |sj(t0)| \u2264 \u03c1j , where \u03b4 > 0. Consider the buck type DC-DC converter which is shown in Fig. 2. Vin is a DC input voltage source, S is a controlled switch, D is a diode, Vo is sensed output voltage, and L,C,R are the inductance, capacitance, load resistance, respectively. The buck type DC-DC converters are used in applications where the required output voltage is smaller than the input voltage. If the switching frequency for S is sufficiently high, the dynamic of DC-DC converters can be described by an average state space model [24]. Based on the average state space model [24], the dynamic equation for the buck converter is: i\u0307L = 1 L (uVin \u2212 Vo), V\u0307o = 1 C (iL \u2212 Vo R ), (47) where u is the control input and u \u2208 [0, 1]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000716_detc2018-85848-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000716_detc2018-85848-Figure6-1.png", "caption": "Figure 6 GE jet engine bracket: (a) conventional design and (b) optimized design for AM", "texts": [ " This study is preliminary research to present the DFAM ontology and its approach to design rule creation rather than providing a complete DFAM system that is easy for a designer to use. To demonstrate the DFAM ontology, we used design guidelines of metal PBF, specifically from a combination of Ti6Al4V material and EOS M280 machine, to build the DFAM ontology. A case study was carried out with the DFAM ontology and the General Electric (GE) jet engine bracket example. The original jet engine bracket in Figure 6 (a) is designed to support the weight of the engine during handling. The bracket may be used periodically but stays attached to the engine at all times. Since the bracket is designed for conventional manufacturing, the bracket is not fully 7 Copyright \u00a9 2018 ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org on 11/06/2018 Terms of Use: http://www.asme.org/about-asme/terms-of-use optimized to trade performance off against weight [31]. By taking advantage of AM technology, the optimized engine bracket design in Figure 6 (b) is introduced to save weight significantly while keeping the original part envelope and performance based on load conditions. [32]. This design came in first place in a GE global 3D printing challenge. Our case study examines the manufacturability of the new design. The design was for powder bed fusion for Ti6Al4V. Minimum wall thickness was provided by GE as 1.27 mm (0.050 in). The geometry parameters of the bracket related to manufacturing features are listed with corresponding values in Table 4", " Based on the normal vector and the unit vector of Z axis that can be represented as =1k, the angle of tilt from the Z axis was calculated as shown in Figure 9. Since the angle is negative, the tilted wall has downskin area which means overhang feature. In step 4, design features were indicated as overhang features that require supports. Equation (6) was used to detect overhang features that require supports based on overhang length and downskin angle for wall features only. Wall features in the DFAM ontology contain two critical features which are horizontal and tilted walls of bracket infill structure as shown in Figure 6 (b). The horizontal wall in the bracket infill region has a 27mm overhanging length with 0\u02da downskin angle. The tilted wall has a 35\u00b0 downskin angle and 16mm overhanging length. According to equation (6) that contains design criteria related to downskin angle and overhang length, these two wall features were recognized as overhang features that require support structure as shown in Figure 10. As a result of the manufacturability analysis, the bracket can be manufacturable with the EOS_M280 machine if the diameters of the hole or pin are changed to meet the clearance requirement and support structures are added in the bracket infill region" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000514_j.mechmachtheory.2019.103667-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000514_j.mechmachtheory.2019.103667-Figure5-1.png", "caption": "Fig. 5. Coordinate systems for tooth contact analysis of face-gear drives.", "texts": [ " Similarly, distance OlOm between the origins of coordinate systems Sl and Sm is given by OlOm = rps sin(\u03b3 + \u03b3 ) + A2 (25) All errors of alignment to be compensated are considered in Eqs. (20)\u2013(23). The equation of meshing is determined as fs2(us, \u03b8s,\u03c8s) = ( \u2202r2 \u2202us \u00d7 \u2202r2 \u2202\u03b8s ) \u00b7 \u2202r2 \u2202\u03c8s = 0 (26) The modified surface of the face-gear with errors of alignment compensated during generation is determined in three- parameter form by vector equation r (us, \u03b8 s, \u03c8 s) and equation of meshing f = 0. 2 s2 Fig. 5 shows the coordinate systems used for tooth contact analysis (TCA) and illustrates the different errors of alignment to be considered: (i) \u03b3 as the shaft angle error, (ii) E as the shortest distance between the axis of rotation of pinion and face gear (axes z1 and z2, respectively), (iii) A1 as the axial displacement of the pinion (not represented in Fig. 5), that is considered in the distance between Oj and Ok, and (iv) A2 as the axial displacement of the face-gear (not represented in the figure either), that is considered in the distance between Ol and Om. Now, the axes of face-gear and pinion will no longer intersect each other but are crossed having E as the shortest distance between axes. Movable coordinate systems S1 and S2 are rigidly connected to the pinion and the face gear. The distance between the origins of coordinate systems Sj and Sk, that includes the axial displacement of the pinion as mentioned above, is given by OjOk = rp2 sin(\u03b3 + \u03b3 ) + rp1 tan(\u03b3 + \u03b3 ) + A1 (27) Similarly, the distance between the origins of coordinate systems Sl and Sm also includes the axial displacement of the face gear and is determined as OlOm = rp1 sin(\u03b3 + \u03b3 ) + A2 (28) The applied tooth contact analysis (TCA) algorithm is based on the rigid body hypothesis of contact of mating surfaces" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000131_acs.2988-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000131_acs.2988-Figure7-1.png", "caption": "FIGURE 7 Switched RLC circuit", "texts": [ " The basis function vectors S1(Z1), S21(Z1), and S22(Z1) contain 23, 27, and 25 nodes, with the centers \ud835\udf071, \ud835\udf0721, and \ud835\udf0722, evenly spaced in [ \u22123, 3] \u00d7 [ \u22124, 4] \u00d7 [ \u22124, 4], [ \u22124, 4] \u00d7 [ \u22122, 2] \u00d7 [ \u22121, 1] \u00d7 [ \u22123, 3] \u00d7 [ \u22123, 3], [ \u22124, 4] \u00d7 [ \u22122, 2] \u00d7 [ \u22122, 2] \u00d7 [ \u22121, 1] \u00d7 [ \u22123, 3], and widths \ud835\udf021 = 2.5, \ud835\udf0221 = \ud835\udf0222 = 2, respectively. The simulation results of Example 1 are shown in Figures 2 to 6. As shown in Figures 2 to 5, all signals in the resulting closed-loop systems (57) and (59) are bounded under a class of switching signals with ADT \ud835\udf0fa = 8 > ln 6.16\u22150.263. The switching signal is depicted in Figure 6. Example 2. The switched RLC system,26 depicted in Figure 7, has been widely employed to perform low-frequency signal processing in integrated circuits. The dynamic equation is given by .x1 = 1 L X2, (61a) .x2 = u\ud835\udf0e(t) \u2212 1 C\ud835\udf0e(t) \u2212 R L x2, (61b) \ud835\udc66 = x1, (61c) where N = 2, L = 1, R = 2, and C1 = 2, C2 = 5. The design procedure is similar that of Example 1. The switched observer design parameters are chosen as l11 = 1, l21 = 2, and l12 = l22 = 2. Furthermore, Q1 = 4I and Q2 = 6I, P1 = [ 6 \u22122 \u22122 4 ] , P2 = [ 4.5 \u22123 \u22123 5.25 ] . The initial values are then set as [x1(0), x2(0)]T = [0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000674_iros40897.2019.8968528-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000674_iros40897.2019.8968528-Figure2-1.png", "caption": "FIGURE 2. Rotating antenna array showing pulses.", "texts": [ " Similarly, the signal of antenna m = \u22121 received by target is s\u22121(t) = Aq(t \u2212 to/2)e j ( 2\u03c0 ( f+f \u22121d ) (t\u2212to/2) ) where Doppler at target from signal of antennam = \u22121 is f \u22121d = (vt\u2212va)/\u03bb = f td \u2212 f a d . At the target, the sum of both signals is received. s+1(t)+ s\u22121(t) = Aq(t \u2212 to/2)e j ( 2\u03c0 ( f+f +1d ) (t\u2212to/2) ) +Aq(t \u2212 to/2)e j ( 2\u03c0 ( f+f \u22121d ) (t\u2212to/2) ) . (3) Let the origin of the array labelled O, P1 and P2 is the wave incident on the sensor after reflection from the target for pulse l = 0 and l = 1 respectively. Let \u03b8o be the angle of incidence for first pulse i.e. l = 0, as illustrated in FIGURE 2. The antenna element m = +1 moves from a point A to point B at the time of second pulse i.e. l = 1 as the antenna undergoes a rotation of angle \u03b8r = \u03c9aT . The wave front travels a shorter distance to reach the sensor m = +1 with angle of incidence \u03b81. Simple trigonometry shows that the difference in path lengths is d (sin \u03b81 \u2212 sin \u03b8o)/2. As a consequence, a Doppler shift is introduced in the echo due to the antenna motion. The same reasoning applies to sensor m = \u22121 with opposite direction of antenna motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000828_j.amc.2015.05.111-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000828_j.amc.2015.05.111-Figure2-1.png", "caption": "Fig. 2. The airframe dynamics model.", "texts": [ " The tip loss and lead-lag motions of blades have been ignored. (3) The tail rotor thrust which balances the antitorque from main rotor can be adjusted by the model automatically and the antitorque from tail rotor has been ignored. In addition, the aerodynamic drag coefficient is considered a constant. 2.2. Interpretation of coordinate system (1) The inertial frame OXY Z: this coordinate system is attached to the earth, which is coincident with body-fixed frame at initial time, its original point is 0.5 m higher than the ground level (shown in Fig. 2). (2) Body-fixed frame O1X1Y1Z1: it is attached to the helicopter body and its origin O1 is located at the midpoint between two rear wheel centers (shown in Fig. 2). 2.3. Force on airframe In order to reflect the forces acting on the airframe, according to simplified result, a 6-DOF airframe model was proposed, as shown in Fig. 2. The six degrees of freedom are: pitching, rolling, yawing of the airframe and translations of the airframe along the three axes of body-fixed frame. The nomenclature in Fig. 2 can be seen in Table 1. 2.4. Aerodynamical forces According to the simplification of model, the aerodynamical forces acting on airframe are main rotor thrust TS, the antitorque from main rotor MK , tail rotor thrust TTR, the aerodynamic drag QF , aerodynamic drag torque Mx and Mz. (1) Main rotor thrust T S Using blade parameters in Table 2, basing on blade element theory, iterative computation formula of main rotor thrust TS and induced velocity \u03c5i are [23]\u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 TS = \u03c1 2 (R ) 2 bcRa \u239b \u239d\u03b8t \u2212 \u03c5i + \u03c5C R 2 \u239e \u23a0 (a) \u03c5i = R a\u03c3 16 [ \u22121 \u2212 8\u03c0\u03c5C abc + \u221a( 1 \u2212 8\u03c0\u03c5C abc )2 + 32\u03b8t a\u03c3 ] (b) (1) where \u03c5C is upward velocity (which is negative when it is downward) and \u03b8t is the pitch angle at the tip", " (2) (2) The antitorque MK In the same way, the antitorque MK can be given by MK = \u03c1 2 bcR2(R )2 \u23a1 \u23a3a \u239b \u239d\u03b8t \u2212 \u03c5i + \u03c5C R 2 \u239e \u23a0\u03c5i + \u03c5C R + cd 4 \u23a4 \u23a6. (3) The meaning of symbols in Eq. (3) can be seen in Table 2. (3) The aerodynamic drags The aerodynamic drags acting on the airfrmam can be simplified to aerodynamic drag QF which is through the mass center and aerodynamic drag torque Mx, Mz. QF = (QFx, QFy, QFz) T = \u2212kS(x\u03072 C, y\u03072 C, z\u03072 C )T , (4) Mx = \u2212kMx\u03b1\u0307, (5) Mz = \u2212kMz\u03b3\u0307 . (6) The meaning of symbols in Eqs. (4)\u2013(6) can be seen in chapter 2.5 and Table 2. 2.5. Geometrical relationship analysis If the time represented in Fig. 2 is t, in a simulation step of t ( t = 0.008s), the infinitesimal displacement of origin O1 relative to body-fixed frame at time t is (x, y, z), the infinitesimal rotation angle of the airframe is (\u03b1, \u03b2, \u03b3 ), transformation matrix representing the orientation of the airframe from time t to t + t will be expressed with matrix A, where for simplicity c\u03b1 and s\u03b1 denote cos \u03b1 and sin \u03b1 respectively, A = \u23a1 \u23a3 c\u03b2c\u03b3 s\u03b3 \u2212s\u03b2c\u03b3 s\u03b2s\u03b1 \u2212 c\u03b2s\u03b3 c\u03b1 c\u03b3 c\u03b1 c\u03b2s\u03b1 + s\u03b2s\u03b3 c\u03b1 s\u03b2c\u03b1 + c\u03b2s\u03b3 s\u03b1 \u2212c\u03b3 s\u03b1 c\u03b2c\u03b1 \u2212 s\u03b2s\u03b3 s\u03b1 \u23a4 \u23a6. (7) When helicopter flies in a simulation step, \u03b1 \u2192 0, sin \u03b1 \u2248 \u03b1, cos \u03b1 \u2248 1, sine and cosine functions of \u03b2 and \u03b3 have the same expressions, matrix A is simplified as A = [ 1 \u03b3 \u2212\u03b2 \u2212\u03b3 1 \u03b1 \u03b2 \u2212\u03b1 1 ] ", " (20) where (x, y, z) denote infinitesimal displacements of original point O1 relative to body-fixed frame at time t and (\u03b1, \u03b2, \u03b3 ) also denote infinitesimal rotation angles of the body-fixed frame from time t to t + t . (2) Kinetic energy of helicopter As the airframe has a mass symmetry plane O1X1Y1, product of inertia Ixz and Iyz all equal zero (other parameters are shown in Table 2). So, when infinitesimal displacements and rotation angles happened, according to Eq. (14), kinetic energy has the following expression: T = 1 2 M ( x\u03072 c + y\u03072 c + z\u03072 c ) + 1 2 (Ixx\u03b1\u0307 2 + Iyy\u03b2\u0307 2 + Izz\u03b3\u0307 2) \u2212 Ixy\u03b1\u0307\u03b2\u0307. (21) (3) Potential energy of helicopter As shown in Fig. 2, at time t, if mass center SC is taken as zero point of geopotential energy and the point located at spring\u2019s original length is taken as zero point of elastic potential energy, potential energy U will equal to the sum of geopotential energy and elastic potential energy as follows: U = U1 + U2. (22) There into, geopotential energy can be given by U1 = Mg sin \u03b8 (x \u2212 L5\u03b3 ) + Mg cos \u03b8 cos \u03bb(y + L2\u03b3 ) + Mg cos \u03b8 sin \u03bb(z \u2212 L2\u03b2 + L5\u03b1). (22.a) Elastic potential energy can be given by U2 = 1 2 3\u2211 i=1 Kix(xi0 + xi) 2 + 1 2 3\u2211 i=1 Kiy(yi0 + yi) 2 + 1 2 3\u2211 i=1 Kiz(zi0 + zi) 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001347_j.ymssp.2014.06.016-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001347_j.ymssp.2014.06.016-Figure16-1.png", "caption": "Fig. 16. Test trajectory for mobile platform.", "texts": [ " Mobile platform validation while moving on desired trajectories (verification of algorithm performance). 3. Algorithm debugging and control system adjustments (R and Q matrices, PWM). Checking procedure of the system performance was executed by measuring deviation between real position of the platform and point 0 of the Cartesian coordinate system 0xy located on the desired trajectory. For each trajectory, the authors established several places where the measurements were taken. The \u201csine\u201d test trajectory, with check-points, is presented (Fig. 16). In Fig. 17 (left) the authors demonstrated result for the last point of the obtained trajectory when the mobile platform moved with 0.17 m/s. In Fig. 17 (right) are presented results for the same check-point, but speed of the platform was 0.31 m/s (90% of max. speed). The test results for both applied speeds confirm sufficiently good performance of the proposed method to control strongly nonlinear object. During the test, the mobile platform reached the final destination with demanding repeatability" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001569_s40846-015-0064-1-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001569_s40846-015-0064-1-Figure6-1.png", "caption": "Fig. 6 False diameter due to aqueous humor in annulus region. a Contact region with presence of aqueous humor, and b example of false diameter image", "texts": [ " The estimated Young\u2019s modulus obtained here is less than the average of previous research, which can be explained by two factors. First, the inclined surface induced a non-circular contact area when the sensor head indented the porcine liver. Most of the contact regions were an ellipse. ImageJ captured the maximum edge in the elliptic region by a circle (see Fig. 4). Therefore, the calculated area was larger than the actual area for a given load, resulting in an underestimation of the Young\u2019s modulus. Second, the aqueous humor on the surface affected the estimation. As shown in Fig. 6, the liquid inside the fresh porcine liver was squeezed out during the indentation process. The aqueous humor adhered to the surface of the sensor head, forming an annular contact region (r0, Fig. 6a). This region is difficult to distinguish and remove from the image. The contrast between the real contact area (rreal) and the annular region was not as clear as that of the annular edge. Thus, the measurement utilized a false contact area. A false diameter is shown in Fig. 6b. The false contact area is larger than the real contact area, leading to underestimation of the Young\u2019s modulus. The aqueous humor effect was not obvious when indenting the PDMS standard or boiled porcine liver because of no or less liquid on the surface. The error caused by the aqueous humor should be removed to obtain the real contact area. Further, the color contrast can also be used to distinguish the annular area. For example, the annular region in Fig. 6b is red because of the blood and the contact region shows the real tissue\u2019s color due to a lack of blood inside. If the width of the red annular area is subtracted from the whole contact area, then the estimated Young\u2019s modulus will be closer to the actual value. The results of the PDMS tests are shown in Fig. 7. The goodness of fit between the experimental data and the simulation data is high (the error is less than 5 %). The estimated Young\u2019s modulus of the PDMS standard is Application of Video-Assisted Tactile Sensor and Finite Element Simulation for Estimating\u2026 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001482_1.4932889-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001482_1.4932889-Figure1-1.png", "caption": "FIGURE 1. The change in the scanning strategy of layers FIGURE 2. Loading pattern of the sample:", "texts": [ " The changes in dimension and form of activated powder particles were registered by the metallographic microscope -RV. Activated and non-activated powders were used for making cubic samples with 10-mm side by VARISKAF100M selective laser sintering machine. The change in ultimate compression strength and in the porosity was considered according to the technological conditions of sintering. Intervals of technological conditions and their critical levels are provided in Table 1. Sintering was carried out in argon shielding atmosphere. The scanning strategy of a sintered layer by laser beam is depicted in Fig. 1. The ultimate compression strength was determined by the INSTRON 5966 desktop universal double-column test system. Two following cases were to be considered to determine the ultimate compression strength: under loading, sintered layers of samples are located in horizontal and vertical planes relative to the surface of the piston. The pattern of loading is shown in Fig. 2. The porosity of samples was determined by weighting as per ISO 2738. RESULTS AND DISCUSSION The mechanical activation causes the increase in bulk density of the powder up to 40% (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002248_j.ijrefrig.2020.01.028-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002248_j.ijrefrig.2020.01.028-Figure1-1.png", "caption": "Fig. 1. (a) Cutaway view of the proposed prototype. (b) A pictorial view of the proposed prototype.", "texts": [ " The effect of gas stiffness shift and esonance frequency change at different input powers were meaured and compared for different gases. Three efficiency evaluation arameters as isentropic, volumetric, and motor efficiencies were efined and measured for the compressor at different operational onditions. The effect of input power on the pressure-volume diarams of different gases was generated to observe the effect of inut powers on the resonance frequencies and overall performance f the compressor. . Linear compressor .1. Design and operation Fig. 1 (a) and (b) presents the cross-sections of the proposed inear compressor. The motor is composed of two circular wound oils around the central axis. The coils are encapsulated in the hree-legged structure as shown in Fig. 1 (a). These coils are con- ected in series, but the direction of the current is opposite to ne another. Therefore, the opposite direction of magnetic flux is roduced in the side legs and the same direction of magnetic flux s generated in the middle leg. As these windings are energized ith the single-phase AC current, a corresponding alternating agnetic flux density is produced in the stator. Consequently, the ower edges of the three-legged structure become temporarily agnetized based on the current magnitude and direction of the lternating current", " uring the positive cycle of alternating current, the interaction etween the stator and mover poles causes a linear thrust force in ne direction. Whereas, the negative cycle of current produces a inear thrust force in the other direction. Therefore, a bi-directional hrust force is generated in the mover due to the alternating curent in the windings. Additionally, an inner core magnetic material alled the back iron provides the return path of the magnetic flux oop in the stator assembly. The overall stator and mover assembly re connected together with three support plates labelled as the ront, back and middle support plate as shown in Fig. 1 (b). The inear motor is placed between the front and middle support late. The mechanical natural frequency of the mover is defined ith the help of eight helical compression springs. Force due to he compressed springs is always opposing the armature\u2019s reciproating motion. The front plate contains the cylinder in which the iston reciprocates. The cylinder is closed with a valve plate on hich the suction and discharge valves are mounted. When the oving assembly goes backward, the suction valve opens and fills he cylinder with the low pressure and temperature refrigerant" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001677_s12206-015-0718-5-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001677_s12206-015-0718-5-Figure3-1.png", "caption": "Fig. 3. The dynamic model of the drive system.", "texts": [ " Dynamic model of the system To consider the contact deformations, the bearings at both ends of the screw and the contact between the screw and the nut are expressed by six stiffness and damping coefficients respectively. Fig. 1 is the bearing\u2019s contact model, and the contact model between the screw and the nut is the same as Fig. 1, but the parameters are different. The contact model of the worktable and guide way is shown in Fig. 2. All parameters are listed in the nomenclature. In the system, the screw rotates at a constant angular veloc- ity \u03a9 . The coordinate origin is at the left end of the screw. The coordinate system and the dynamic model of the system are shown in Fig. 3. The worktable translation ( )s t in x direction includes axial translation produced by the screw rotation, axial one mlj caused by torsional deformation of the screw, ( ),du tlq from axial deformation of the screw and xq from local vibration of the worktable, and is expressed as: ( ) ( ),d m d xs t u t qlq lf lq= + + + . (1) Kinetic energy of the system, including the kinetic energy of the screw sT and the kinetic energy of the worktable mT , is expressed as: s mT T T= + . (2) The system potential energy, including the screw deformation energy Us and elastic contact energy between the system elements, such as the contact energy between the worktable and the screw Un, the contact energy of the bearings Ub and the contact energy between the worktable and the guide way Um, is expressed as: " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001243_s00170-015-6914-8-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001243_s00170-015-6914-8-Figure2-1.png", "caption": "Fig. 2 External forces applied on the cantilever and particle, which have been modeled by the presented relations.", "texts": [ " \u03b8 :: \u00bc \u03b8 :2 tan\u03b8 \u00f03\u00de \u03b3\u0308 \u00bc 2\u03b3\u03072 sin \u03b3 2 cos \u03b3 2 cos2 \u03b3 2 \u2212sin2 \u03b3 2 \u00f04\u00de Dynamic of nanoparticle manipulation has been investigated in recent works for both spherical and cylindrical particles; in this part, a previously introduced model has been used. In the dynamic manipulation of nanoparticles using the AFM robot, two groups of forces are of interest, which include (1) internal forces consisting of Fx; Fy; Fz;M \u03b8X ;M \u03b8Y ;W ;V , and (2) external forces consisting of FX,FY,FZ. The internal forces are actually the reactions of external forces, which have been simulated by means of spring constants; and in order to express their relationships and the deformations produced in the cantilever, relation (5) has been used (Fig. 2). The reaction forces of Fv,Fu,Fw, which are applied from the probe to the particle, have been employed in the investigation of motion modes [13, 23]. FX FY FZ 2 4 3 5 \u00bc KX 0 0 0 K\u03b8Y H 0 KY 0 K\u03b8X H 0 0 0 KZ 0 0 2 6664 3 7775 \u03b5X \u03b5Y \u03b5Z \u03b8X \u03b8Y 2 66664 3 77775 \u00f05\u00de In the above relation, FX ,FY, FZ are the applied forces; KX ; KY ; KZ ; K\u03b8X ; K\u03b8Y are the spring constants; \u03b5X ,\u03b5Y, \u03b5Z are the displacements; and \u03b8X ,\u03b8Y are the angular displacements. Knowing the internal forces, the Newton-Euler relations can (a) Front view (b) Side view \u23ad \u23ac \u23ab \u23a9 \u23a8 \u23a7 = p p Y X probeP \u23ad \u23ac \u23ab \u23a9 \u23a8 \u23a7 = t t Y X tipt Cantilever\u2019s bending \u03b3 ) 2 sin() 2 sin(2 \u03b3\u03b3H ) 2 cos() 2 sin(2 \u03b3\u03b3H 2 \u03b3 2 \u03b3 Cantilever\u2019s twisting \u03b8cosH \u03b8sinH Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure6-1.png", "caption": "Fig. 6. Dynamic model of gear transmission system in gearbox.", "texts": [ "69 Hz \u2014 ous influencing factors, such as time-varying meshing stiffness, meshing damping, tooth backlash, error of tooth surface, and friction of the tooth surface. A bending-torsional-axis coupling dynamic model for the transmission system of gearbox is established in this section to calculate the inner dynamic excitation and analyze the impact of different factors on dynamic excitation. The eccentricity error affects the high-speed gear pair much more than low-speed pair. Thus, only first stage gear pair (highest-speed stage) with eccentricity error situation is considered in this study. Fig. 6 shows the dynamic model of the gearbox transmission system. In this model, lumped mass and lumped rotational inertia are used to simulate the helical gear at all stages. Nonmass rigid body is used to simulate transmission shaft. Spring and damper are used to simulate the elastic support of the bearing. The gear teeth engagement is simulated by a pair of springs and dampers. The torsional vibration of the shaft between gears 2 and 3, gears 4 and 5, and gears 6 and 7 is considered. The equivalent mass, damping, and stiffness of each component in the model are determined by the structure size of the gear, shaft, and bearing. When defining angular displacement, the direction of rotation of each gear under the action of input torque is the positive direction of torsional vibration. As shown in Fig. 6, gear 1 is in the input, and gear 8 is in the output. kxi, kyi, and kzi are the support stiffness values and cxi, cyi, czi are the damping values of each bearing. The input torque is Tin, and the output torque is Tout. kmi and cmi are the meshing stiffness and damping of each gear pair. k23, k45, and k67 are the torsional stiffness values and c23, c45, and c67 are the torsional damping values of each shaft. The generalized displacement array can be expressed as { }T, , ,i i i ix y z \u03b8 , (30) where xi, yi, and zi are the displacement of each gear in X direction, Y direction, and Z direction, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002117_auv.2016.7778652-FigureI-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002117_auv.2016.7778652-FigureI-1.png", "caption": "Fig. I. The equipment layout ofSBPAUV", "texts": [], "surrounding_texts": [ "SBPAUV can behave underwater with the combination of predefmed behavior patterns. The patterned behavior control helps SBPAUV's operators to set up a mission plan easier. Predefmed behavior patterns of SBP AUV are as follows Pattern 1. Following target heading and depth It is pattern to the depth and direction control directed to a certain direction. Commonly used to check the operation of in the tank test. AUV systems started to keep when the depth and heading angle has become in the vicinity of the specified value, or at the set time, the pattern is completed. Pattern 2. Steady turn Steady turning clockwise as viewed from above. In the case of reception of the command from acoustic communication, or at the set time, AUV systems sift the next step. Pattern 3. Spiral ascend / descend Turning clockwise as viewed from above. In the case of achievement of a specified depth or altitude, AUV systems sift the next step. Commonly used after drop the ascending weight or before drop the descending weight. Pattern 4. Moving toward the Way Point at a constant depth When the measured depth is considered to be not correspond with a specified depth, AUV systems control so as to approach the value. By setting the altitude limit, AUV does not approach the bottom of the sea than its value. 978-1-5090-2442-1 /1 6/$31.00 \u00a9201 61EEE Pattern 5. Moving toward the Way Point at a constant altitude When the measured altitude is considered to be not correspond with a specified altitude, AUV systems control so as to approach the value. By setting the depth limit, AUV does not dive than its value. Pattern 6. Moving toward the Way Point until AUV reaches a specified depth or altitude Moving toward the Way Point and, once it become a specified depth or altitude, pattern is completed. Pattern 7. Changing to a specified depth while going to the Way Point While complementing the depth, moving toward the Way Point. Pattern 8. Vertical zig-zag maneuver between the specified depth 82 Moving toward the Way Point, while a reciprocating between the specified depths. III. SEA TRIAL After fmishing the assembling, the land test and the water tank test, we conducted the sea trial for SBP AUV in Suruga Bay from December 7 to 17, 201S. The objective of this trial is the overall performance check of SBPAUV and the on board calibration of devices, which are difficult to be conducted by the land or water tank test. Suruga Bay that has a gentle slope of about Imile and the depth is 80\ufffd90m is suitable adjustment place for AUV. Details of the sea trial are as follows. \\. The adjustment of the weight and the buoyancy ii. The integrated performance check of the acoustic communication system and the acoustic positioning system Ill. The operation check of on board devices IV. The speed - power consumption relation v. The test of each patterned behavior VI. The control gain tuning VII. The overall dive test by combined behavior patterns By these test, performances of SBP AUV proved to be satisfactory. And details of each dives are follows. 1. Cruising test on the sea surface (dive #1) Considering risk of lost the AUV, SBP AUV was performed cruse on the sea surface at the begirming. The setting course is a square of one side about 370m. The setting speed was 1.Sm/s. And we check the behavior of the iridium beacon. The positioning accuracy of the iridium beacon is about IS\ufffd2Sm compared to inferred position. AUV2016 2. Dive test of a constant depth (dive #2-1, #2-2) Next, we were confIrmed of dive at a constant depth. The setting course is a straight line of about 370m one way (dive #2-1) and a square of one side about 370m (dive #2- 2). The speed was l.5m/s and the depth was 20m. First, we conducted a straight line test, SBPAUV completed the mission. The fIrst overshoot was about l.lm, but in subsequent cruising the overshoots ware within the range of pI us or minus 30cm. Moreover, in the square test SBPAUV could control the depth within the range of plus or minus 30cm. 978-1-5090-2442-1 /1 6/$31.00 \u00a9201 61EEE 83 3. Dive test for the control gain tuning(dive #3-1, #3-2) SBP AUV dived for control gain tuning of the longitudinal and lateral motions. For the longitudinal motion the course is the ascend/descend course within the depth range of 10- 60m (dive #3-1). In the dive for lateral motion SBP AUV dived zig-zag course in the depth of 20m (dive #3-2). The speed was changed from l.5 mls to l.Om/s during the dive. In the result of dives, SBPAUV performed stable control at both of dive, therefore we did not change the control gain. 4. Dive test for the compute of the endurance (dive #4) In the next dive that course was a round-trip of the straight line of about llOOm in the depth of 20m. The speed of the AUV was changed in such order as l.5m/s, l.Om/s, O.5m/s during this dive. We measured the consumed power and calculated the endurance with respect to each speed . The appropriate speed is l.Om/s for operation of SBP. Table 2 shows the endurance of SBPAUV is 16h, which mean that SBPAUV is able to dive over the target time 12 hours. TABLE 11. THE SPEED - ENDURANCE RELA nON Speed O.5m/s 1.0m/s 1.5m/s Endurance 24h 16h 9h AUV2016 S. Dive test of a constant altitude (dive #S) SBPAUV usually dives with a constant altitude for SBP. Therefore in this dive, we set altitude to SOm, and the speed changed l.Sm/s to l.Om/s during the dive. At the speed of I.Sm/s, SBP AUV cruised keeping the target altitude within the range of plus or minus 1m. At the speed of l.Om/s, SBPAUV cruised keeping the target altitude within the range of plus or minus SOcm. Speed O.Sm/s 1.0m/s 1.3m/s Turning diameter ISm 20m 20m 978-1-5090-2442-1 /1 6/$31.00 \u00a9201 61EEE 84 7. Dive test for spiral ascend I descend and acoustic communication test (dive #7) In dive #7, we checked spiral ascend/descend pattern. We set the target depth to 60m, the turning diameter to 30m and the speed to l.Om/s and O.Sm/s. When the speed shifted to O.Sm/s, we also checked the acoustic communication. Communication situation is satisfactory, position update was also successful. 8. Dive test by combined behavior patterns At the end of the sea trial, we tried the overall dive test on the assumption that we perform the real operation of SBPAUV.SBPAUV completed the mISSIOn safely, performances of SBPAUV proved to be satisfactory. ACKNOWLEDGMENT This work is supported by Council for Science, Technology and Innovation (CSTI), Cross-ministerial Strategic Innovation Promotion Program (SIP), \"Next-generation Technology for Ocean Resources Exploration\" (funding agency: Japan Agency for Marine-Earth Science and Technology, JAMSTEC). REFERENCE [I] Kangsoo Kim, Kenkichi Tamura \"The Zipangu of the Sea Project Overview: Focusing on the R&D for Simultaneous Deployment and Operation of Multiple AUVs,\" Proc. of the Offshore Technology Conference Asia 201 6, Kuala Lumpur, Mar. 201 6, OTC-26702-MS. AUV2016" ] }, { "image_filename": "designv11_22_0001222_j.msea.2014.06.107-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001222_j.msea.2014.06.107-Figure4-1.png", "caption": "Fig. 4. Hysteresis stress\u2013strain curve for the material in the state of cyclic plasticity during bearing operation (\u03c4: shear stress, \u03b3: shear strain, \u03c41;max: maximum principal shear stress).", "texts": [ " When the material starts to deform plastically, a residual stress is generated. In a two dimensional contact configuration, the residual stress can reduce \u03c3x, resulting in a reduction of \u03c41 [2,9]. However, since \u03c4xz remains constant, \u03c41 can only be reduced down to \u03c4xz maximum, 0:25p0. Therefore, \u03c41 could only increase up to 0:25p0 in the presence of a residual stress. Considering cyclic plasticity when the material state is between elastic and plastic shakedown [10], the stress\u2013strain curve for each cycle would form a closed loop as shown in Fig. 4. The maximum principal shear stress (\u03c41;max) would then be in the range of 0:25p0r\u03c41;maxr0:30p0 depending on how much residual stress has been developed. The responsible strain for the plastic deformation would be between 0 and AB as some of the deformation is reversed. In this study, it is assumed that \u03c41;max \u00bc 0:25p0 and that the strain responsible for the plastic deformation per cycle is \u0394\u03b3C \u00bc 1 2AB. 100Cr6 steel was used in this study. The alloy composition is shown in Table 1. The specimens were heat-treated in the form of 15 15 25 mm3 blocks", " The block was mechanically polished with silicon carbide papers, 6 \u03bcm and 1 \u03bcm diamond suspension, and colloidal silica for X-ray diffraction analysis. Data were obtained with a 2\u03b8 range of 35\u20131251 with a 0.041-step and 5 s-dwell time per step employing a Philips PW1820 diffractometer using Cu K\u03b1 radiation at 40 kV and 40 mA. By Rietveld refinement with Fullprof version 0.50 software, the measured amount of austenite was 15.370.3 vol% with (goodness of fit)2\u00bc3.94. The blocks were finally cut into cylinders of 3.43 mmdiameter and 5.00\u20135.21 mm-length. The repetitive push test in this study was set up to reproduce the response in Fig. 4 with uniaxial stress. The test consisted of uniaxial compressive stress cycles as shown in Fig. 5. During the test, a specimen is loaded up to pmax during the first 30 s, and compressive stress cycles were applied for 5 105 cycles with a frequency of 15 Hz. Finally, unloading took place for 1 s. During the stress cycling, a minimum stress of 0.04 GPa was imposed in order to prevent the specimen from slipping from the stage. Note that 15 Hz is the maximum frequency that the employed tester manages to apply the stress stably. pmax was varied from 1.0 to 3.5 GPa with a 0.5 GPa-step. The maximum resolved shear stress (\u03c4rss;max) for each case can be obtained from \u03c4rss;max \u00bc \u00f01=M\u00depmax, where M is the Taylor factor equal to 2.9 for randomly oriented body-centred cubic metals [11]. Considering Fig. 4, \u03c4rss;max can be considered to be equivalent to the maximum principal shear stress (\u03c41;max). Then, pmax can be related to the contact pressure (p0) in the bearing tests. According to Section 2, \u03c4rss;max \u00bc 1 M pmax \u00bc \u03c41;max \u00bc 0:25p0; \u00f01\u00de \u2018 p0 \u00bc 1 0:25M pmax: \u00f02\u00de For the tested pmax, the corresponding \u03c4rss;max and p0 values are listed in Table 2. The test employed a Mayes 100 kN mechanical tester with the setup schematically shown in Fig. 6. The cross section of the lower stage was a 76 mm-sided square and that of the upper stage was a 30 mm-diameter circle", " Strain gages (KFG-1N-120-C1-11 from Kyowa) were attached to the side of the samples with CC-33A strain gage cement (cyano-acrylate base). The gage was the smallest available, containing a 1 0.65 mm2 grid on a 4.2 1.4 mm2 base. The sample was loaded by moving the upper stage while the lower stage was kept stationary. Despite the high pressure, no barreling was observed after the test, possibly due to the high hardness of the sample. Examples of the obtained data from the strain gage are given in Fig. 7a. These plotted data sets represent the normal equivalents of the shear hysteresis curve in Fig. 4. To determine \u0394\u03b3C , 2\u0394\u03f5C are firstly obtained from each curve as shown in Fig. 7b; a horizontal line on pmean \u00bc \u00f0pmax\u00fe0:04\u00de=2 (pmean and pmax in GPa) was drawn, and the difference of the intersecting strain values was taken. These \u0394\u03f5C values are defined to be \u0394\u03f5C;gage. Then, 2\u0394\u03f5C;gage is plotted with respect to the number of cycles (N) as shown in Fig. 8 for pmax \u00bc 3:0 GPa. In all cases, \u0394\u03f5C;gage showed a decay with respect to N as shown in Fig. 8 as a result of strain hardening with time. \u0394\u03f5C;gage approached an asymptotic value after a certain N, which would correspond to the strain responsible for the plastic deformation (\u0394\u03f5C) for a long run bearing" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002672_ccdc49329.2020.9164017-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002672_ccdc49329.2020.9164017-Figure2-1.png", "caption": "Fig 2. D-H link coordinate system of left arm.", "texts": [ " The dulcimer music robot in this paper consists of a mechanical arm, a wrist striking device, a support base and a dulcimer. The three-dimensional model is shown in Fig. 1. According to the manipulator configuration of dulcimer This work is supported by the Planning Fund of the Department of Social Sciences of the Ministry of Education under Grant 16YJAZH080. music robot, the kinematic model of the robot arm is established by D-H parameter method [10]. Because the structures of the left and right arms are same, the left arm is modeled in this paper, the D-H link coordinate system is shown in Fig. 2, and the link parameter of left arm is shown in Table 1. 4769978-1-7281-5855-6/20/$31.00 c\u00a92020 IEEE Authorized licensed use limited to: Carleton University. Downloaded on September 21,2020 at 05:58:00 UTC from IEEE Xplore. Restrictions apply. Where, 1\u2212ia is the length of the link 1\u2212i , 1\u03b1 \u2212i is the angle of the link 1\u2212i , id is the distance between link i and link 1\u2212i , \u03b8i is the angle of joint i . Positive kinematics is used to solve the end position and posture of the robot arm under the given angle of each joint and the link parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003718_j.mechmachtheory.2021.104425-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003718_j.mechmachtheory.2021.104425-Figure3-1.png", "caption": "Fig. 3. (a) Schematic for machining worm W g by a planar grinding wheel according to D. Brown Co.; (b) Diagram for installation position of planar grinding wheel.", "texts": [ " To solve these questions, a rigorous proof of whether a worm gear tooth surface with constant helix parameter ground by a planar grinding wheel is an involute surface is presented in this section using the uniqueness theorem for surfaces [33] . The mapping and rigid body transformation relations between these two surfaces are also studied. The proof in this section complements the machining principle of the machining method proposed by D. Brown Co. for the involute worm ground with the planar grinding wheel. In this section, the symbol W g indicates the cylindrical worm machined by the planar grinding wheel and the symbol S w represents the worm tooth surface enveloped by the family of planes. As shown in Fig. 3 (a), during processing the worm tooth surface, a unit orthogonal static frame \u03c3o1 { O 1 ; i o1 , j o1 , k o1 } is fixed on the worm roughcast. The unit basic vector k o1 is along the axis of the worm roughcast. The original point O 1 is located at the middle point of the worm thread length. The plane O 1 \u2212 j o1 k o1 is parallel to the horizontal plane. A unit orthogonal rotating frame \u03c31 { O 1 ; i 1 , j 1 , k 1 } is set to denote the current position of the worm roughcast whose rotational angle is represented by \u03d5. The shortest distance between the points O 1 and O od is called the operating center distance a d when machining the worm roughcast. In the actual processing, the worm roughcast needs to be rotated around its axis k o1 while the grinding wheel makes the translational motion along the negative direction of the axis \u2212\u2212\u2212\u2212\u2192 O od O d . This is because the worm studied in this section is right-handed. Thereupon, in the current position shown in Fig. 3 (a), the length of the straight-line section | \u2212\u2212\u2212\u2212\u2192 O od O d | is | \u2212\u2212\u2212\u2212\u2192 O od O d | = p\u03d5 where the symbol p represents the helix parameter of the worm tooth surface S w . Two translational frames \u03c3od { O od ; i od , j od , k od } and \u03c3d { O d ; i d , j d , k d } are used to represent the installation position of the planar grinding wheel. The plane O 1 \u2212 j od k od is the horizontal plane. The angle between the vectors k od and \u2212\u2212\u2212\u2212\u2192 O od O d is 90 \u25e6 \u2212 \u03b21 where the angle \u03b21 is the helix angle of the worm tooth surface. The rotational angle between the frames \u03c3od and \u03c3d along the axis j d is the angle \u03b1 which is called the normal pressure angle of the worm. According to Fig. 3 (b), the equation of the generating plane can be represented in \u03c3d as ( rd ( u, v ) ) d = u i d + v j d , (26) where u and v are two variables of the generating plane. From Eq. (26) , the unit normal vector, nd , of the generating plane can be obtained in \u03c3d as ( nd ) d = \u2212 k d . (27) Employing the rotational transformation matrix, the equation of the generating plane in Eq. (26) and its unit normal vector nd in Eq. (27) can be represented in \u03c3od as ( rd ) od = R [ j od , \u03b1 ] ( rd ) d = [ cos \u03b1 0 sin \u03b1 0 1 0 \u2212 sin \u03b1 0 cos \u03b1 ] ( rd ) d = u cos \u03b1 i od + v j od \u2212 u sin \u03b1 k od , (28) ( nd ) od = R [ j od , \u03b1 ] ( nd ) d = [ cos \u03b1 0 sin \u03b1 0 1 0 \u2212 sin \u03b1 0 cos \u03b1 ] ( nd ) d = \u2212 sin \u03b1 i od \u2212 cos \u03b1 k od " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure6.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0003481_s00170-021-07005-6-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003481_s00170-021-07005-6-Figure5-1.png", "caption": "Fig. 5 Temperature distribution during the deposition of a 1st; b 2nd; c 6th; and d 10th layer deposition", "texts": [ " When heating pulse approaches nearest to the thermocouples in each layer deposition, temperature sharply increases. The temperature peak becomes higher at successive layers. This trend is attributed to the high heating injection by laser but relatively slow thermal dissipation at the substrate. The thermal energy accumulates in short time and increases the temperature in successive layers. This statement can be confirmed by the temperature evolution at different times during 10 layers deposition of Ti6Al4V, shown in Fig. 5. The highest temperature is located at the laser pulse center and the temperature becomes smaller away from the laser pulse. There is a heat affected zone where the temperature is relatively high than remote zone. This heat affected zone moves with the laser pulse. When this heat affected zone is closest to the neighboring thermocouples, measured temperature experiences a peak. Absorbed thermal energy increases substrate temperature gradually. Figure 6a shows the temperature evolution in 1st to 10th layers at the center of the left straight wall", ", 1650 \u00b0C), third to tenth peak temperature are lower than the melting temperature. This means, the first layer is affected by ten thermal cycles, but can only be re-melted by the second track. For the only temperature peak of last layer deposition, highest temperature peak of 2250 \u00b0C was obtained. In addition, the temperature contours present elongated shape behind the heating beam and compressed shape ahead of heating beam, which is caused by the rapid scanning motion of laser beam. The highly transient and spatially non-uniform temperature distribution, shown in Fig. 5 and Fig. 6a, are responsible for the generation of the stress and strain fields. Figure 6b represents the cooling rate of four distinct locations (1st, 2nd, 6th, and 10th layer) and shows that the maximum cooling rate for each monitoring point occurs exactly as laser scans on that layer. As deposited layers increase, the maximum cooling rate for each individual layer gradually reduces due to the increased temperature of previous layers and the substrate. The temperature gradient (especially Z direction component) between two consecutive monitoring points maximizes at 1st~2nd layer and gradually reduces with more layer deposited" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002914_978-3-030-60330-4-Figure2.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002914_978-3-030-60330-4-Figure2.4-1.png", "caption": "Fig. 2.4 Dispersion curves of the discrete model", "texts": [ " The physical meaning of the Brillouin zone boundaries consists in that they show the following values of the wave vectors or the electron quasipulses, in which the electron wave cannot propagate in a solid [39]. Next, we will analyze the dispersion properties of the medium in the first Brillouin zone and on its boundary depending on the values of the microstructure parameters. Like in the solid-state physics, each normal lattice vibration can be associated with a certain type of quasiparticle\u2014phonon [39, 40]. The considered system has a longitudinal acoustic (LA) phonon, a transverse acoustic (TA) one, and a rotational optical (RO) phonon (Fig. 2.4) [41, 42]. We pass to the polar coordinate system q1 = q cos \u03b8 , q2 = q sin \u03b8 , in Eq. (2.18), where q is the wave vector module and the angle \u03b8 indicates the direction of the plane wave propagation with respect to x-axis in the direct lattice. In particular, in the case of propagation of the plane waves, when q2 \u2261 0 and, hence, d12 \u2261 d13 \u2261 d21 \u2261 d31 \u2261 0, Eq. (2.14) is substantially simplified since the longitudinal phonons become independent in it: ( 2 \u2212 d11 K0 )(( 2 \u2212 d22 K0 )( 2 \u2212 d33 K0 ) \u2212 d23 K0 d32 K0 ) = 0, (2", "Consequently, by varying the microstructure parameters (see Eq. (2.4)), it is possible to specify certain dispersion properties of the phonon crystal [23, 26]. Let us perform analysis of solutions of the dispersion Eq. (2.18) for the following values of the microstructure parameters: d/a = 0.1, K1/K0 = 0.5, K2/K0 = 0.3. The dispersion curves calculated along directions \u03b8 = 30\u00b0 (G\u2013K), \u03b8 = 30\u25e6 (G\u2013M) and along the boundary of the Brillouin zone (K\u2013M) are shown in the dimensionless coordinates (qa, ), where = \u03c9 / \u03c90 and \u03c90 = \u221a K0/M , in Fig. 2.4. From Fig. 2.4, it is visible that in the G-M-direction the frequencies of all three phonons increasemonotonically, when thewavenumber grows, up to the boundary of the Brillouin zone. In the G-K-direction, the frequency of the longitudinal phonons has a local maximum \u2248 3.63 located at the point qa = 2 ( \u03c0 \u2212 arctg ( 3 \u221a 7 )) . In the interval 2 ( \u03c0 \u2212 arctg ( 3 \u221a 7 )) < qa < 4\u03c0/3, the group velocity of rotational phonons is negative: vgr = d\u03c9/dq < 0. This area is called a backward-wave region [39]. Usually, a field of the negative group velocity exists for optical phonons in lattices with a complex structure, when more than one particle is present in the Bravais lattice [43]", " The frequency of the transverse phonons has the maximal value \u2248 3.36 at the point K. The rotational (optical) mode has two threshold frequencies: the minimum (0) \u2248 3.45 and maximum \u2248 4.18 ones. In the frequency range 0 \u2264 \u2264 3.36, the system has LA- and TA-modes. In the interval 3.36 < < 3.45 there is only a longitudinal mode and for frequencies 3.45 \u2264 \u2264 3.93 there are longitudinal and rotational modes. And, finally, in the high-frequency range 3.93 < \u2264 4.18, only the rotational mode is present in the system (Fig. 2.4). Figure 2.5 shows maps of equal frequencies for longitudinal, transverse, and rotational phonons (for LA-mode = 1.0, 1.5, 2.1, 2.7, 3.0, 3.3, 3.6, 3.8, for TAmode = 0.7, 1.0, 1.5, 2.1, 2.7, 3.0, and for RO-mode = 3.65, 3.75, 3.85, 3.95, 4.05, 4.15) [44]. The horizontal axis represents the projection qx , of the wave vector, and along the vertical axis\u2014qy . The boundaries of the first Brillouin zone are indicated by a dashed line. Figure 2.5 shows that lines of equal frequencies are circles for small values of the wavenumber" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003009_j.mechmachtheory.2020.104223-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003009_j.mechmachtheory.2020.104223-Figure1-1.png", "caption": "Fig. 1. A CDPR for UAVs as a lightweight gripper.", "texts": [ " Additionally, generating rotational trajectories that must respect kinematic constraints as well as consistency conditions is still an open issue. In this paper, an approach is proposed for planning dynamic rotational trajectories that can pass through the actuation singularities of a 3-DOF rotational CDPR, using the unit quaternion. The 3-DOF rotational CDPR is designed as a lightweight gripper with high-speed rotational capacity, suspended under unmanned aerial vehicles (UAVs), as illustrated in Fig. 1 . A rigid link is placed between the UAV and the moving platform with a spherical joint on the moving platform, restricting the translations of the moving platform. Four cables are applied to redundantly actuate the CDPR, the design of which is a compromise between maintaining its simple lightweight structure and optimizing its rotational performance. The structure of this paper is as follows. In Section 2 , the dynamics represented by the unit quaternion are derived. The actuation singularity loci of the mechanism are analyzed and the consistency conditions of the dynamic equation are then obtained in Section 3 ", " + \u03b2m \u03b3m = 0 , (29) where \u03b3p (p = 1 , ., m ) represents the pth element of \u03b3 . Suppose \u03b2k = 0 , Eqs. (26) and (29) can be rewritten as \u03b21 \u03b2k M 1 + . . . + \u03b2m \u03b2k M m = M k , (30) \u03b21 \u03b2k \u03b31 + . . . + \u03b2m \u03b2k \u03b3m = \u03b3k , (31) where \u03b21 \u03b2k , . . . , \u03b2m \u03b2k , the same as \u03b1p in Eqs. (24) and (25) , denote the linear combination coefficients. Letting \u03b2p \u03b2k = \u03b1p , Eq. (31) is equivalent to Eq. (25) . Additionally, as stated in [22] , the rank r of M usually becomes m \u2212 1 . Therefore, for the 3-DOF mechanism shown in Fig. 1 , the rank r becomes two at an actuation singular orientation and only one consistency condition is required to be satisfied in order to pass through singular orientations, written as \u03be \u00b7 \u03b3 = 0 . (32) where \u03be is the column vector which span the null space of M at a singular orientation. In this section, dynamic point-to-point rotational trajectories using the unit quaternions are planned to reach a series of target orientations. In the presence of singular orientations, a transition trajectory is designed to pass through singularities by meeting the consistency conditions obtained in Section 3 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure12-1.png", "caption": "Fig. 12. Constraint of analysis model.", "texts": [ " Only the important parts of the entire test rig, i.e. top frame main and side assembly, base frame assembly, test bearing set, adaptor, and insert plate were included in the analysis because the other parts have less influence on the result. Furthermore, as the number and size of bolts were carefully selected through relative guidelines [11], no bolts were included in the analysis model. Instead, all bolt-connected positions of the test rig were fixed together in all directions by the RBE2 method, as shown in Fig. 12. The final analysis model contained 321127 mesh elements\u2014124970 shell elements for the top frame main and side assembly; 152476 solid elements for the bearings, adaptor, and insert plate; 43175 shell elements for the base frame assembly; and 506 rigid elements for the boltconnected positions\u2014which have a hexagonal or tetragonal shape according to the location, and the mesh size was in the range of 50-60 mm. All hydraulic cylinder loadings were applied to the corresponding positions of the test rig, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003734_aer.2021.53-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003734_aer.2021.53-Figure1-1.png", "caption": "Figure 1. Controlled variables and reference axes at initial time.", "texts": [ " Furthermore, we demonstrate the applicability of our backstepping solution for two fixed-wing aircraft having very different configurations and specific properties, also in presence of noise and parametric uncertainties. 3.0 FIXED-WING AIRCRAFT MODEL Fixed-wing aircraft dynamics are defined by a six-degree-of-freedom model. Three sets of differential equations describe the forces and moments acting on the aircraft and its orientation with respect to a reference system(29). The force equation in generic Centre of Gravity (CG)-centred body axes (XB, YB, ZB) (Fig. 1) is expressed as mV\u0307 = F \u2212 \u03c9 \u00d7 mV, \u00b7 \u00b7 \u00b7 (1) with m aircraft mass, V = (u, v, w)T linear velocity vector and \u03c9 = (p, q, r)T angular velocity vector, all expressed in body axes. The vector F = (Fx, Fy, Fz)T represents the sum along XB, YB and ZB of all forces acting on the aircraft centre of mass: aerodynamics forces, engine thrust and gravity force. The moment equation has a similar structure: I\u03c9\u0307 = M \u2212 \u03c9 \u00d7 I\u03c9, \u00b7 \u00b7 \u00b7 (2) where I is the body-axes inertia matrix. The vector M = (M , L, N)T contains the sum of the moments about XB, YB and ZB generated by aerodynamic forces and engine thrust", " The scalar ub is the external controller of the global system; each subsystem represented by the state \u03bel (l = 1, ..., k \u2212 1) is controlled by the virtual control input \u03bel+1. Equations (1)\u2013(3), as such, cannot assume the structure of (4). The cascade form is not respected since F and M are function of the states V and \u03c9, of the aerodynamic angles and of control actions. Nevertheless, under the assumptions described below, it is possible to convert the equations of motion into a suitable form for a limited number of aircraft states: angleof-attack \u03b1, sideslip angle \u03b2 and stability-axes roll rate ps (Fig. 1). The aim is to design a controller such that \u03b1= \u03b1ref , ps = pref s and \u03b2 = 0. Control over angle-of-attack and roll rate is essential to determine, respectively, the longitudinal behaviour and the flight direction. A null sideslip angle is desired in cruise flight to achieve symmetric flight and reduce aerodynamic drag. The ability of an aircraft to cancel out sideslip angle perturbations is a sign of its lateral\u2013directional static stability. Stability axes are a particular type of body axes where XS lies along the projection of the initial V on the aircraft plane of symmetry, ZS is positive from the upper to the lower side of the wing aerofoil and YS completes the right-handed reference frame (Fig. 1). XS and XB are separated by the angle-of-attack, and a single rotation of magnitude \u03b1 about YS \u2261 YB is sufficient to align body axes with stability axes. Such rotation allows the definition of the angular velocities in stability axes \u03c9s = (ps, qs, rs)T as \u03c9s = \u23a1 \u23a2\u23a3 cos \u03b1 0 sin \u03b1 0 1 0 \u2212 sin \u03b1 0 cos \u03b1 \u23a4 \u23a5\u23a6 \u03c9 = Rsb\u03c9. \u00b7 \u00b7 \u00b7 (5) The dynamics considered for the control design are obtained from the force equation written in wind axes; the complete derivation is available in Ref. (2). Wind axes are defined as follows: XW is aligned with the airspeed direction, YW is orthogonal to XW oriented from left to right with respect to the centre-of-mass trajectory and ZW lies in the plane of symmetry of the aircraft, directed from the upper to the lower wing aerofoil surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000790_rcar.2018.8621743-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000790_rcar.2018.8621743-Figure1-1.png", "caption": "Fig. 1. mecanum wheel Fig. 2. The coordinate of wheel", "texts": [ " The attraction force is Fatt= grad [Uatt(q)] = \u2212Ka |q \u2212 qgoal| (2) The repulsion potential function is expressed as Urep(q) = { 1 2 Kr [ 1 d(q,qobs) \u2212 1 d0 ]2 d \u2264 d0 0 d > d0 (3) where Kr is a repulsive constant factor, d(q, qobs) = |q \u2212 qobs|, d0 is the crucial distance. The repulsion force is Frep = \u2212grad [Urep(q)] = { Kr [ 1 d(q,qobs) \u2212 1 d0 ] \u2202d(q,qobs) d(q,qobs) d \u2264 d0 0 d > d0 (4) The total potential function is U = Uatt + Urep. (5) The total force is F = Fatt + Frep. (6) III. KINEMATICS MODEL The Fig. 1 shows the mecanum wheel which includes the hub and the passive rollers. From the Fig. 2, the following equation can be calculated[ v\u2032ix v\u2032iy ] = [ 0 sin\u03b1i r cos\u03b1i ][ \u03c9i vir ] = Ki1 [ \u03c9i vir ] (7) where all parameters are in the wheel coordinate, v\u2032ix, v\u2032iy, \u03c9i are the wheel speed and angular velocity. vir represents the norm velocity of the wheel. r represents the radius of the wheel. It is necessary to translate v\u2032ix, v\u2032iy in the x\u2032oy\u2032 coordinate to vix, viy in the xoy coordinate. From the Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002886_jae-209506-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002886_jae-209506-Figure14-1.png", "caption": "Fig. 14. Distribution of magnitude of magnetic flux density for load state (right magnitude of first harmonic of induction).", "texts": [ " Table\u00a011 presents the core losses caused by the dominant higher harmonics generated by the inverter calculated with use circuit analytical method. Table\u00a012 gives a comparison of the motor losses calculation results obtained using the circuit analytical model and field \u2013 circuit model with the measurement results when the motor is powered from the inverter. The influence of higher harmonics generated by the inverter is very low at the fundamental harmonic frequency of the supply voltage of 350\u00a0Hz and causes a rise in the core losses below 10% (7.5%). As can be seen from Fig.\u00a014, the flux density level in the motor for 350\u00a0Hz power supply is quite low (0.7\u00a0T) compared to 10\u00a0Hz or 20\u00a0Hz operation (1.8\u00a0T). This is because for this frequency the motor is already working in the field weakening region because of the limited voltage at DC input. Low saturation of the magnetic circuit additionally increases losses from higher harmonics. Figures\u00a015 and\u00a016 show selected harmonics of magnetic flux density. These are harmonics produced by the rotor slots (32 slots from here 16 harmonics)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002673_j.matpr.2020.07.229-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002673_j.matpr.2020.07.229-Figure1-1.png", "caption": "Fig. 1. Geometric configurat", "texts": [ " Comparison of natural frequencies and their displacements have been done. A 3D solid model was created in SolidWorks as a rectangular cross section which can also be incorporated in an aerofoil profile. The outer dimension of the cross-section is 64 mm X 24 mm with the beam thickness as 4 mm. The length of the beam was taken as 680 mm. This beam has been considered as a cantilever beam and the above solid model was imported in ANSYS and modal analysis was conducted for obtaining the values of natural frequencies for different mode shapes as shown in Fig. 1. The rectangular beam has been analysed with Aluminum 2024, E-glass fibre and Carbon fibre. The fibre orientation of glass fibre and carbon fibre is considered to be zero. The material properties have been given below in Table 1: Any engineering system requires the correct definition of its motion. To define the motion, the force needs to be balanced using inertia force, damping force and spring force. The equation of motion for a system (Eqn.1) with n degrees of freedom is given by, M\u20acx\u00fe C _x\u00fe Kx \u00bc F \u00f01\u00de Where M, C and K are the n n mass, damping and stiffness matrices" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003547_j.jmapro.2021.04.060-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003547_j.jmapro.2021.04.060-Figure8-1.png", "caption": "Fig. 8. The sketch of subarea-sequential and subarea-parallel mode.", "texts": [ " The detailed algorithm description of the sequential scanning pattern is given in the supporting information (Fig. S5(a)). mode. The scanning length is dependent on the OPW in the subarea scanning mode. A long scanning length leads to large deformations along the scanning direction, and a short scanning length always induces heat concentrated in a small area [12]. The classical rectangular monolayer is employed to determine the OPWs corresponding to the subarea-sequential and subarea-parallel modes (Fig. 8). As shown in Fig. 9, the total length of the rectangular is equally divided into 1\u20136 parts respectively. The proposed matrix-randomized scanning mode is derived from the LaserProFusion technique of the EOS company [27]. This technique utilizes nearly one million array heat sources (laser, diode) to sinter the entire powder layer instantaneously. As shown in Fig. 10, For the conventional single laser technics, the laser scanning process of each layer will mostly take several seconds or more. In contrast, the matrix-laser system can obviously reduce the building time for each layer into millisecond level regardless of the complexity of the scanning profile" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001845_cp.2016.0182-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001845_cp.2016.0182-Figure2-1.png", "caption": "Figure 2: Geometric interpretation of the MTPA condition", "texts": [ " Having for the MTPA condition, the correspondent is determined by (7) In order to solve the optimality condition online using Newton Method, a geometrical interpretation of the MTPA optimization, firstly introduced in [11], can be used. Then, the optimality condition can be solved using Newton method. Let be the vector tangent to the torque curve at each point. For an , denotes the vector that is perpendicular to the current circle at each point on it. The idea is based on the fact that, for each point on : implies that the point is on the MTPA curve, i.e. it is a tangent point between and . implies that the point lies on the right side of the MTPA curve. implies that the point lies on the left side of the MTPA curve. Fig. 2 illustrates this principle. It is important to highlight that the scalar product between two vectors is equal to the cosine of the angle between them. One may note that The gradient of T is defined by: (8) Where and are unitary vectors in the direction of the dand q-axis respectively. Let , \u03b1,\u03b2 . For each MTPA point, it must satisfy , which implies that (9) By choosing and , equation (9) holds. Then, a vector which is tangent to T at each point, is given by (10) For the current circle , the gradient vector, eliminating constant factors, is: (11) Thus, the cosine of the angle between and can be found through the scalar product of these two vectors" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003009_j.mechmachtheory.2020.104223-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003009_j.mechmachtheory.2020.104223-Figure2-1.png", "caption": "Fig. 2. Architecture of the 3-DOF rotational CDPR.", "texts": [ " Four cables are applied to redundantly actuate the CDPR, the design of which is a compromise between maintaining its simple lightweight structure and optimizing its rotational performance. The structure of this paper is as follows. In Section 2 , the dynamics represented by the unit quaternion are derived. The actuation singularity loci of the mechanism are analyzed and the consistency conditions of the dynamic equation are then obtained in Section 3 . In Section 4 , dynamic rotational trajectories based on the unit quaternion interpolation are designed to pass through singular orientations. Finally, Section 5 draws the conclusion. As illustrated in Fig. 2 , a fixed reference frame F 0 with its origin O on the static platform is defined. The z-axis of the fixed reference frame is pointing upward, i.e., opposite to the direction of gravity. Point A j ( j = 1 , . . . , 4) is fixed and corresponds to the location where the cable j exits its pulley. A moving reference frame F 1 is defined on the moving platform with its origin o located at the centre of the spherical joint. Point P represents the geometric centre of the moving platform. B j ( j = 1 , . . . , 4) is the attachment point of the moving platform and each cable. According to the vector closure principle, the inverse kinematic equations can be written as \u03c1 j = a j + Qb j + h , (1) where \u03c1 j is the vector from point A j to point B j along the jth cable, a j is the vector from point A j to origin O, b j represents the vector oB j in the reference F 1 . h is the vector of the rigid link, namely the vector from O to o, as shown in Fig. 2 . Q is the rotation matrix from the fixed frame to the moving frame. The effective cable length \u03c1 j is given by \u03c1 j = \u221a (a j + Qb j + h ) T (a j + Qb j + h ) . (2) e j = \u03c1 j /\u03c1 j is defined as the unit vector in the direction of the jth cable, oriented from the static platform to the moving platform. Free-body diagram of the moving platform is shown in Fig. 3 . Then the dynamic model is built using the NewtonEuler approach (all frictions are ignored), which yields 4 \u2211 j=1 f j + \u03ba + m g = 0 , (3) 4 \u2211 j=1 (Qb j ) \u00d7 f j + Qp \u00d7 (m g ) = Q (I \u02d9 \u03c9 + \u03c9 \u00d7 I \u03c9 ) , (4) where m is the mass of the moving platform", " After retrogression, the seventh degree polynomial consisting of eight unknown parameters can be determined. In this section, an example is provided to verify the rotational trajectory planning scheme proposed in Section 4 . Singularity loci are first determined and a trajectory segment that connects two consecutive target orientations is designed. Then series of target orientations are simulated to show the feasibility of the scheme for passing through singular orientations of the CDPR. Dimensions of the 3-DOF rotational CDPR shown in Fig. 2 are as follows. The radius of the static platform is r 0 = 120 mm and the radius of the moving platform is r 1 = 50 mm . The length of the rigid link is h = 110 mm . The mass of the moving platform without load is 44 g and the mass point vector in the coordinate system F 1 is [0 , 0 , \u221210] mm . The offset distance between the cable attachment points on the moving platform and the rotation centre in the z 1 direction represented by symbol e is \u22127 mm. Due to the fact that the unit quaternions are not intuitive for the representation of orientations, target orientations are first described using the tilt-and-torsion (T & T) angle convention" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001221_esda2014-20001-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001221_esda2014-20001-Figure2-1.png", "caption": "FIGURE 2.FILLETS OF A MARINE DIESEL ENGINE CRANKSHAFT. COURTESY OF BM SERVICES S.R.O.", "texts": [ " The principal design of the in-situ laser cladding device which is proposed in this article is subject to the International Patent application PCT/LV2013/000006 of 18.07.2013 \u2013 Apparatus and method for in-situ repair and renovation of crankshaft journal surfaces by means of laser cladding [15]. The damaged crankshaft journal surface is renovated by fitting the laser cladding nozzle positioning and guidance device directly on the crankshaft journal fillets. These fillets as a rule are not damaged or worn and thus conserve the original manufacturers\u2019 crankshaft dimensions (see Fig.2). The internal fillets can therefore be used as a reference base, enabling the laser cladding nozzle guidance platform to be positioned using the device\u2019s specially designed guide-ways. Drawings of the \u0160koda/\u010cKD type \u201c6-27.5 A2L\u201d medium-sized marine diesel engine crankshaft fillets in Fig. 3 are provided as an illustrative example. In this case, the diameter of the crankshaft main bearing is 210 mm with internal fillet (see Y in Fig.3) radius of 11 mm, and the crankpin journal is 190 mm with internal fillet radius of 12 mm (see X in Fig.2). Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 2014 by ASME The essence of the research is a device for the repair and renovation of crankshaft journal surfaces. Its technical description is provided below to highlight the complexity and challenges as well as the clear benefits of this endeavour. The device comprises two guide-ways and two opposite guide-ways for positioning the device on crankshafts fillets and two frame parts, each of which is fixed to the respective guideway" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001435_s10439-014-1185-3-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001435_s10439-014-1185-3-Figure1-1.png", "caption": "FIGURE 1. Schematic of testing apparatus for in vivo testing of the impulsively end loaded human upper extremity. Each subject lay on a padded table with left hand positioned on a force transducer (F). The subject was asked to concentrate on monitoring EMG biofeedback from his/her elbow extensor muscle activity provided on a display screen (S) and maintaining it at a certain level of muscular effort. A weight (W) of 23 kgf was then released by a remote trigger after a random delay to strike the end of the lever-arm (B) in order to apply an impulsive force to the wrist, thereby causing elbow flexion and shoulder adduction (the end of the lever-arm changing from B to B\u00a2). Alpha (a) and theta (h) represent the initial shoulder extension and elbow flexion angles, respectively.", "texts": [ " We then measured subject\u2019s resting and maximum voluntary pre-contraction (MVC) electromyographic (EMG) levels of the triceps brachii (long and lateral head), biceps brachii (short head), the anterior deltoid, pectoralis major and serratus anterior muscles during elbow and shoulder flexion, and extension and ab- and adduction by pulling up or pushing down on an handle attached to a vertical cable in series with an uniaxial force transducer (TLL-500, Transducer Techniques, Temecula, CA, USA) used for measuring maximum arm strengths. The several configurations for the elbow and shoulder muscle strength tests are described in the Appendix, Fig. A. Next, each subject was asked to lie prone on a table with the left arm vertical and wrist positioned on a 6 axis force transducer (MC3A-1000, AMTI, Newton, MA, USA) mounted at one end of a 76 mm 9 152 mm 9 2,032 mm hollow aluminum beam having a rectangular cross-sectional shape and wall thickness 6.35 mm (Fig. 1). The beam was pivoted at its midpoint about a fulcrum formed from a pair of collinear 1(Protocol #HUM00052983). needle bearings mounted on an axle in the horizontal plane. An Optotrak Certus camera (Northern Digital, Inc., Waterloo, Canada) was used to measure the displacements of 15 infrared-emitting optoelectronic markers (shown as dots in Fig. 1). A marker or marker triad were taped on the left side of the force transducer (markers #1\u20133), on the left aspect of the hand (#4), on the most caudal-lateral point on the radial styloid (#5), a triad on the left lateral aspect of the middle of the forearm segment (#6\u20138), at the most caudal point on lateral epicondyle (#9), a triad at the left lateral aspect of the mid upper arm segment (#10\u201312), at the center of rotation of the glenohumeral joint (#13), on the left lateral aspect of the base of the neck (#14), and over the T1 spinous process. When ready, the subject was asked to hold the heel of his/her hand \u2018\u2018lightly\u2019\u2019 in contact with the force transducer and to concentrate on monitoring EMG biofeedback from his/her lateral head of the triceps muscle on a display screen maintaining triceps activity either at rest, or 25, 50, and 75% MVC values from the main agonist muscle. A weight of 23 kgf was released (shown as \u2018\u2018W\u2019\u2019 in Fig. 1) from a height of 720 mm to impact the top surface of the other end of the beam, thence applying an upward impulsive force to the wrist via the force transducer, thereby causing elbow flexion, shoulder extension/adduction and trunk extension. The subject was instructed \u2018\u2018not to intervene\u2019\u2019 before, during and after the weight drop. For example, if the trial was conducted at 50% triceps MVC, the subject was instructed to contract the target muscle steadily at 50% of MVC during the test. Three trials at least were conducted at each of the three levels of muscle activation, and these were presented in randomized order" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001588_s12206-015-1118-6-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001588_s12206-015-1118-6-Figure1-1.png", "caption": "Fig. 1. The lead screw system: (a) the components of the lead screw system; (b) the test setup.", "texts": [ " The frequency separation between the torsion and axial modes was controlled by the location of the nut. The frequency of the axial modes was increased when the nut moved on to the middle of the lead screw, whereas that of the torsion mode was not changed. This study also identified the friction curve by measuring the friction torque when the nut slide along the lead screw. Furthermore, to verify the vibration mode of the frequency of friction noise, the hammering test were performed with respect to the location of the nut. As shown in Fig. 1(a). The lead screw system consists of the lead screw and the nut. By the rotation of the lead screw controlled by the DC motor, the nut moves back and forth as in Fig. 1(b). Fig. 2 demonstrates a system diagram of the experimental equipment used in this study. A axial load was applied to the nut by a pulling weight. The axial load was controlled by the change of the weight. A torque sensor was attached between the motor and lead screw to measure friction *Corresponding author. Tel.: +82 41 521 9272, Fax.: +82 41 554 6516 E-mail address: keysun@kongju.ac.kr \u2020 Recommended by Associate Editor Junhong Park \u00a9 KSME & Springer 2015 torque, where a linear position sensor was used to measure the travel position of the nut" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000959_s00502-014-0272-3-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000959_s00502-014-0272-3-Figure8-1.png", "caption": "Fig. 8. Cut-views of the designs of the three machines ((a), (b), (c)) for equal magnet mass mMa = 1.9 kg (simulated with JMAG Designer 12.1)", "texts": [ " Due to the same electromagnetic thrust \u03c4 and the bigger air-gap area A\u03b4 of the AFM, the currents of the AFM could be reduced to get the same torque. This led to higher efficiencies in OP1 and OP2 for the AFM and equal efficiencies in OP3. But the used magnet mass of the AFM (mM,AFM = 2.3 kg) was higher than that of the RFM (mM,RFM = 1.9 kg). For a fair comparison the magnet masses should be equal. Therefore the magnet height of the axial flux machines were reduced to hM,AFM = 5 mm to get the same magnet mass (Fig. 8). Due to the reduction of magnet height the calculated efficiencies in OP1 and OP2 of the axial flux machines are reduced, while the efficiency in OP3 is increased due to the lower needed field weakening current. Table 3 summarizes the efficiencies and the active masses of the three designed machines. The calculation of the efficiencies includes the ohmic losses and the iron losses in stator and rotor. The friction losses and additional losses like eddy-current losses in the magnets were neglected. The calculated efficiencies of the designed machines at OP1 and OP2 are nearly equal. At OP3 the efficiencies of the axial flux machines are higher due to the reduced magnet height. Hence, the axial flux machines show a slight advantage over the radial flux machine. 28 heft 1.2015 \u00a9 Springer Verlag Wien e&i elektrotechnik und informationstechnik Table 3. Calculated efficiencies and active masses of the three designed machines of Fig. 8 Fig. 8 Variant Efficiency OP1 Efficiency OP2 Efficiency OP3 Active mass (a) AFIR 86.6 % 93.8 % 90.4 % 32.0 kg (b) AFIS 86.5 % 93.9 % 89.7 % 32.5 kg (c) RFM 86.4 % 93.9 % 88.3 % 32.6 kg Fig. 9. Calculated flux density distribution of a half section of a quarter of the AFIS-machine at no-load (Max. value: 1.9 T, Min. value: 0 T) Figure 9 shows the flux density distribution of the AFIS-machine at no load. All three machines are designed for an air-cooling system to keep the complexity of the driving system as low as possible" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001853_1350650116632018-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001853_1350650116632018-Figure2-1.png", "caption": "Figure 2. Progression of contact and emergence of new asperity.", "texts": [ " They used straight line profiles to connect the peak ordinates and named their asperity as n-point asperity where n is the indicator of number of peak ordinates of which the asperity is composed of. Figure 1 illustrates the basic difference in definition of these asperities. As the n-point asperity model developed by Hariri et al.9 is the foundation of the present work, it is necessary to outline in brief the salient features of this model in order to set a scene for the formulation of the present problem. As can be seen in Figure 2(a) and (b) in n-point asperity model, asperity curvature as well as height changes with the progression of contact. Also with decrease in separation, the previous asperity gets merged into a new asperity with higher number of n-points. Thus as compared to conventional Greenwood and Williamson model,12 the npoint asperity model represents the rough surfaces in more realistic form as it considers the variation in form of asperities in vertical direction (asperity height direction) as well as horizontal direction (asperity spacing direction)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001083_1.4027926-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001083_1.4027926-Figure8-1.png", "caption": "Fig. 8 Faults in 6206 ball bearing of aero-engine rotor experimental rig", "texts": [], "surrounding_texts": [ "4.1 Faults Produced on Ball Bearings. To study the casing responses caused by ball bearing faults, in this study, faults are produced on ball bearings using wire-electrode cutting technology. These faults are produced on the inner ring, outer ring, and ball of the compressor experimental rig\u2019s 6214 ball bearing and the aero-engine experimental rig\u2019s 6205 ball bearing, which are shown in Figs.7 and 8. The ball bearing dimensions are listed in Table 1. The ball bearing faults\u2019 characteristics frequencies can be computed as follows:\n(1) Outer race\nfoc \u00bc Z 2 1 d D cos a fR (1)\n(2) Inner race\nfic \u00bc Z 2 1\u00fe d D cos a fR (2)\n(3) Rolling ball\nfbc \u00bc Z\n2\nD d 1 d D cos a\n2 \" #\nfR (3)\n(4) Cage\nfc \u00bc 1 2 1 d D cos a fR\n(4)\n4.2 Faults Experiments of Ball Bearings. The bearing faults experiments are carried out. The test-site photo of the compression rotor rig is shown in Fig. 9, and the measurement points\u2019 explanation is listed in Table 2; the test-site photo of the aeroengine rotor rig is shown in Fig. 10, and the measurement points\u2019 explanation is listed in Table 3.\nVibration signals are collected by means of the USB9234 data acquisition card of the NI Company, the 4805 type ICP acceleration sensors of B&K Company are used to pick up the acceleration signals, and the eddy current sensors are used to measure the rotating speeds. The sampling frequency is 10.24 kHz.\n4.3 Wavelet Envelope Analysis for Ball Bearing Fault\n4.3.1 Basic Principles of Wavelet Envelope Analysis. In this study, a signal analysis to determine ball bearing faults is carried out by means of a wavelet envelope spectrum analysis. Chen [2] provides a reference for the detailed process of this algorithm. The essence of ball bearing fault diagnosis based on the wavelet packet is to take advantage of its bandpass filter characteristics, and to decompose the signals using appropriate wavelet functions so as to obtain an appropriate resonance frequency band. Then, by means of envelope demodulation, low frequency envelope signals that only contain the fault characteristic information are obtained. Its spectrum is the wavelet envelope spectrum, in which the fault characteristic frequencies of the ball bearings can be found out.\n4.3.2 Wavelet Envelope Spectrum Analysis of Ball Bearing Fault Signals\n4.3.2.1 Experiment Analysis Based on Compressor Rotor Experimental Rig\n(1) Feature extraction for inner ring faults\nFigures 11\u201314, respectively denote the time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope spectra of the bearing house\u2019s response and the casing\u2019s response to inner ring faults in the 6124 ball bearing. The experimental rotating speed is 1793 rpm (29.88 Hz). The number of balls is 10, the other ball bearing\u2019s parameters are listed in Table 1, and the inner ring characteristic frequency can be calculated by formula (2) as fic\u00bc 5.8974 29.88\u00bc 176.24 Hz. From Fig. 11, the vibration acceleration response value of the casing is larger than that of the bearing housing, which shows that the impulse vibration of the bearing fault is effectively transmitted to the casing. This result is in full agreement with the experimental result for the impulse response. From Fig. 12, it can be found out that there are many resonance peaks in the frequency spectrum, and the signals are very weak in the low frequency segments, the signal\u2019s frequency spectrum from 0 Hz to 500 Hz are shown in Fig. 13, from which, the characteristic frequency of the inner ring fault and its modulation frequency fr can be\nJournal of Dynamic Systems, Measurement, and Control NOVEMBER 2014, Vol. 136 / 061009-3\nDownloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "basically seen. The wavelet envelope spectrum is shown in Fig. 14, by comparison with Fig. 11, the characteristic frequency of the inner ring fault and the modulation frequency fr are more distinctly shown in Fig. 14. Obviously, the inner ring fault features can be all extracted from the acceleration signals of the bearing house and the casing by means of the wavelet envelope spectrum analysis and frequency spectrum. (2) Feature extraction of outer ring faults The time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope\nspectrum of the bearing house responses and casing responses of the 6124 ball bearing outer ring faults are respectively shown in Figs. 15\u201318. The experimental rotating speed is 1826 rpm (30.43 Hz). The number of balls is 11, the other ball bearing\u2019s parameters are listed in Table 1, and the outer ring characteristic frequency can be calculated by formula (1) as foc\u00bc 4.5879 30.43\u00bc 140 Hz. From Fig. 15, the vibration acceleration response value of the casing is larger than that of the bearing housing, which shows that the impulse vibration of the bearing fault is effectively transmitted to the casing. This result is in full agreement with the experimental result for the impulse response. From Fig. 16, it can be found out that there are many resonance peaks in the frequency spectrum, and the\nTable 1 Ball bearing dimensions (in mm)\nType Diameter of inner ring Diameter of outer ring Thick Diameter of ball Pitch diameter\n061009-4 / Vol. 136, NOVEMBER 2014 Transactions of the ASME\nDownloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use", "signal\u2019s spectrum from 0 Hz to 500 Hz is shown in Fig. 17, from which, the characteristic frequency of the outer ring fault can be distinctly seen. The wavelet envelope spectrum is shown in Fig. 18, by comparison with Fig. 17, the characteristic frequency of the outer ring fault are also distinctly shown in Fig. 18, obviously, outer ring fault features can be all effectively extracted from the acceleration signals of the bearing house and casing by means of frequency spectrum and the wavelet envelope spectrum analysis. In addition, no modulation frequency components appear in the frequency spectrum and the wavelet envelope spectrum. (3) Feature extraction for ball faults The time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope spectrum of the bearing house\u2019s response and casing\u2019s response for the 6124 ball bearing\u2019s ball faults are respectively shown in Figs. 19\u201322. The experimental rotating speed is 1827 rpm (30.43 Hz). The number of balls is 11, the other ball bearing\u2019s parameters are listed in Table 1, the ball fault characteristic frequency can be calculated by Formula (3) as fbc\u00bc 5.3919 30.43\u00bc 164.1 Hz, and the cage rotating frequency can be calculated by formula (4) as fc\u00bc 0.4103 30.43\u00bc 12.5 Hz. From Fig. 19, the vibration acceleration response value of the casing is larger than that of the bearing housing, which\nFig. 10 Test-site photo of aero-engine rotor experimental rig\nTable 3 Explanation of the measurement points of the aeroengine rotor experimental rig\nChannel Measurement variable Measurement point position MOD3, CH1 Rotating speed Coupling MOD3, CH2 Acceleration Bearing house MOD4, CH1 Acceleration Casing top MOD4, CH2 Acceleration Casing right MOD4, CH3 Acceleration Casing bottom MOD4, CH4 Acceleration Casing left\nJournal of Dynamic Systems, Measurement, and Control NOVEMBER 2014, Vol. 136 / 061009-5\nDownloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use" ] }, { "image_filename": "designv11_22_0002914_978-3-030-60330-4-Figure6.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002914_978-3-030-60330-4-Figure6.2-1.png", "caption": "Fig. 6.2 Rotational degrees of freedom of the particle", "texts": [ " We consider a cubic lattice consisting of rigid spherical particles (grains) of massM and having a shape of a sphere with diameter d. In the initial state, they are located in the sites of the lattice with a period a (Fig. 6.1). Each particle possesses six degrees of freedom: The center of gravity of a particle with number N = N (i, j, k) can move along the x-, y-, and z-axes (translational degrees of freedom ui, j,k , vi, j,k , and wi, j,k), and the particle itself can rotate around each of these axes (rotational degrees of freedom \u03b8i, j,k , \u03c8i, j,k , and \u03d5i, j,k) (Fig. 6.2). In this case, the kinetic energy of the N th particle is described by the following formula: T = M 2 ( u2t + v2t + w2 t ) + J 2 (\u03d52 t + \u03b82 t + \u03c82 t ), (6.1) where J = 2 5M ( d 2 )2 = 1 10Md2 = 2 5M 3b2 4 = 0.3Mb2 is the moment of inertia of the particle with respect to each axis passing through its center of gravity. \u00a9 Springer Nature Switzerland AG 2021 V. Erofeev et al., Structural Modeling of Metamaterials, Advanced Structured Materials 144, https://doi.org/10.1007/978-3-030-60330-4_6 129 As in themodels discussed in the previous chapters, the space between the particles is an inertia-less elastic medium through which the force and moment effects are transmitted, which are modeled by elastic springs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001219_s11538-014-9991-1-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001219_s11538-014-9991-1-Figure2-1.png", "caption": "Fig. 2 Examples of MT movement caused by motors. a MT reorientation by a minus-end-directed motor, b MT sliding", "texts": [ " One function of many motors is that they are able to walk along MTs (either towards their plus end or minus end), carrying important proteins with them, distributing them to the appropriate locations within cells. Motor proteins can also affect MT organization. They can do so by either (1) helping to align MTs parallel with one another by MT cross linking, and/or (2) by aiding in MT-directed transport (MT sliding). MT alignment occurs when motor proteins are cross-linked to two MTs simultaneously (Ned\u00e9l\u00e9c et al. 1997). As they walk along the MTs simultaneously, they produce pushing and pulling forces that help to reorient the MTs (see Fig. 2a). Some motors can walk long distances along MTs (and are called processive), while others may only walk short distances or not at all (these motors are non-processive). MT sliding occurs when a motor is attached (absorbed) to a non-moving substrate at its cargo domain, where its free legs are able to attach to a MT (Yokota et al. 1995; Gibbons et al. 2001). Since the motor remains stationary, it effectively pushes the MT along its own axis as it walks along it. Such a sliding mechanism is often replicated in in vitro experiments, and is referred to as a gliding assay (Yokota et al. 1995; Gibbons et al. 2001; Vale et al. 1992; Tao et al. 2006) (see Fig. 2b). Sliding speed has been found to vary greatly in such experiments because it depends directly on motor speed. Over the past few decades, both local and non-local models have been proposed to describe how MT patterning occurs in systems composed of motor proteins and MTs. MT sliding is described by simple advection (directed movement), placing these models in the category of transport-type models. Transport-type models are defined as models where the particles of interest are defined by their position in space, time, and velocity (Perthame 2007)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001314_icaccct.2014.7019402-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001314_icaccct.2014.7019402-Figure1-1.png", "caption": "Figure 1. Twin rotor MIMO system", "texts": [ " The TRMS nonlinear model obtained from the Feedback systems operation manual [1] is linearized and is used for H\u221e controller design. The rest of the paper is organized as follows: a description of the system along with the mathematical model is exposed in Section II. Section III describes the classical decoupling technique in TRMS. Section IV and V includes H\u221e Mixed sensitivity controller design. Section VI deals with the simulation results and conclusions in section VII TRMS is a laboratory size helicopter model built to study the complications involved in the actual helicopter. As shown in figure.1, TRMS mechanical unit consists of a beam pivoted on its base in such a way that it can rotate freely both in its horizontal and vertical planes. At both end of a beam, there are two propellers driven by DC motors. The two motors are placed perpendicular to each other. The aerodynamic force is controlled by varying the speed of the motors. Therefore, the control inputs are the supply voltages of the DC motors. The TRMS system has main and tail rotors for generating vertical and horizontal propeller thrust" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure3.3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure3.3-1.png", "caption": "Fig. 3.3 The assumption of a tangential contact between two oriented point pairs can be used to define a relative transformation AT B . Graphic taken from [4]", "texts": [ "3 Results of Dessimoz\u2019 work on gripper pose estimation . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 7 Figure 2.4 Visualization of the three important poses in the workspace of a bin-picking robot . . . . . . . . . . . . . . . 11 Figure 3.1 Regular structure of a 3D mesh acquired by a laser line scanner. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Figure 3.2 Rotation and translation invariant features of a dipole . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Figure 3.3 The assumption of a tangential contact between two oriented point pairs can be used to define a relative transformation ATB . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 Figure 3.4 Pose hypotheses generation using a \u2018birthday attack\u2019like approach. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 Figure 3.5 Scan of piston rods lying on a table (SICK LMS400) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 Figure 3.6 Visualization of the occurring problem when sheet metal parts are scanned from a single point of view ", " Obviously, it is not efficient to exhaustively search through the 6d space of all relative locations. Therefore, only pose hypotheses with a certain surface contact between model and scan are considered. Such a hypothesis can be constructed by assuming a contact between some surface points. More precisely, four given oriented surface points a, c \u2208 A and b, d \u2208 B are sufficient, if a tangential contact between a and b as well as between c and d is assumed. This assumption constrains all degrees of freedom of the relative transformation. As illustrated in Fig. 3.3, the homogeneous 4\u00d74 transformationmatrix AT B can be estimated bymeans of two predefined frames (one coordinate system for the CAD model an one for the scan): AT B(a, b, c, d) = F(a, c)\u22121 F(b, d) (3.1) where the function F(u, v) is a homogeneous 4\u00d74 transformation matrix, representing a coordinate system located between the points u and v of a dipole 3.1 Generic Pose Estimation Using 3D Point Clouds 17 F(u, v) := [ puv \u00d7 nuv\u2016 puv \u00d7 nuv\u2016 puv puv \u00d7 nuv \u00d7 puv\u2016 puv \u00d7 nuv \u00d7 puv\u2016 pu+ pv 2 0 0 0 1 ] (3" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003395_j.matpr.2021.02.158-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003395_j.matpr.2021.02.158-Figure1-1.png", "caption": "Fig. 1. 3D Solid model of the proposed three-wheeler \u2018\u2018battery\u201d vehicle [1].", "texts": [ " In India, delta type chassis are commonly used in the threewheeled vehicles including both \u2018\u2018auto-rickshaw\u201d and \u2018\u2018electric rickshaw\u201d due to the advantages of less turning curvature and rear wheels drive architecture as compare to tadpole type configuration [22]. Delta type configuration has one front wheel and two rear wheels while tadpole type configuration having two front wheels and one rear wheel. The primary objective of this research is to design the chassis of a three-wheeled \u2018\u2018electric\u201d vehicle using SolidWorks software further to compute the inertia relief translational and rotational acceleration of unsupported structure in ANSYS Workbench tool. Fig. 1 shows the 3D Solid model structure of the three-wheeler \u2018\u2018battery\u201d vehicle prepared in the SolidWorks software. Steel grade IS513 having yielding strength of 225 MPa is used for the construction of conventional auto rickshaw chassis [22]. AISI 1005 steel is employed for the body structure of an automobile car [23]. Magnesium alloy based lightweight chassis is developed for four-wheeler vehicle [24]. Mild steel is extensively employed in the existing three-wheeler \u2018\u2018battery\u201d vehicles chassis [14]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002111_1.4035481-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002111_1.4035481-Figure2-1.png", "caption": "Fig. 2 Schematics of the domains used for the femoral implant problem. (a) Orthographic schematic of the idealized design domain, with proximal torsion and fixed distal boundary conditions shown, (b) Isometric sketch of ideal domain, (c) 2D slice of rasterized ideal domain (top right quarter only), (d) Isometric sketch of domain with collar added, and (e) 2D slice of rasterized domain with collar (top right quarter only).", "texts": [ " To reduce nonhomogeneous influences to the optimization by the mesh, a regular mesh is used to describe the prosthetic design Journal of Biomechanical Engineering MARCH 2017, Vol. 139 / 031013-3 Downloaded From: http://biomechanical.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jbendy/936016/ on 01/31/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use and solve the linear elastic problem. The mesh containing the whole domain has 128 128 512 hexahedral elements. The volumes XB (the existing bone) and XP (the prosthetic) need to be rasterized so as to fit this. Figure 2(c) shows a top view of the ideal domain rasterized into voxels\u2014note that the surface P is partially inside XB and partially inside XP. The surface P being contained in elements of both material types is problematic: The calculated stress on P is quite irregular, causing rasterization artifacts in the evaluation the optimization objective. To mitigate this issue, we modify the problem domain slightly to that of Fig. 2(d), by adding a collar (XC) to the interface between the prosthetic and the bone. This allows the surface P to exist entirely in the same material and smooths out the integrand on P relevant for calculation of the objective function. The material properties inside XC are set to be the same as the existing bone and XC XP. 2.3 Shear Stress Objective. The objective function to be minimized is [8] F \u00bc \u00f0 P \u00f0sisi\u00dem dS \u00bc \u00f0 P s2m dS (3) where s is the shear stress vector with norm s and P is the cylinder with radius r1 ranging over z [z2, z3] between the existing bone and prosthetic (shown in Fig. 2). Conceptually, the objective should reduce shear stress concentrations on the bone-prosthetic interface by penalizing high values via the exponent 2m. As m is raised, it is expected that the optimization will distribute the shear stress more evenly throughout the interface. There will be stress singularities at corners such as at the top of the interface surface P [19]. For F to be well-defined, m must be small enough such that the singular integrand is still integrable. If m is too large, then we may still calculate a value numerically, however, the value of the integral may not converge but instead be mesh dependent" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure2-1.png", "caption": "Fig. 2. Coordinate systems of internal meshing gear pair.", "texts": [ " Through giving spatial surface A and its motion law, the other spatial surface B, which is conjugated to surface A, can be obtained. Specially, the surface B is unique due to the limitation of spatial position relationship [2]. The proposed conjugate curve theory considering a variety of contact form of the spatial curves shows that the spatial curve C2 can be got if given a spatial curve C1 and its designated motion law. It is noteworthy that the spatial curve C2 is not unique because of the various contact position relationships. The comparisons of two theories are displayed in Fig. 1. Fig. 2 shows the coordinate systems of internal meshing gear pair, including the fixed coordinate systems S(O-x, y, z), Sp(Op-xp, yp, zp), and the movable coordinate systems S1(O1-x1, y1, z1), S2(O2-x2, y2, z2). \u03c91 and \u03c92 are rotational angle velocities, which are, respectively, corresponding to gears 1 and 2. 1\u03d5 and 2\u03d5 are the rotational angles of gears 1 and 2 under the given motion. There is a relationship 2 21 1i\u03d5 \u03d5= , where i21 is transmission ratio. a is central distance and point P is the common contact point of internal gear pair" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002520_j.matpr.2020.05.438-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002520_j.matpr.2020.05.438-Figure2-1.png", "caption": "Fig. 2. (a) FDM mechanism (b) S", "texts": [ "75 mm) available as a flexible monofilament. This is heated to semi-liquid state. Then, it is extruded through a deposition head (which is numerically controlled) on a fixtureless table in a controlled environment (temperature). The nozzle movement in x-y plane is numerically controlled with respect to base through which complex geometric model can be created preciously. As it finishes deposition the first layer of filament material fabrication on the bed the building platform moves downward (refer Fig. 2) for the next layer thickness deposition. Thereby, single layer is repeatedly deposited at a time on the previous one till the component is manufactured. In general, depending upon the component design and drawing, the support material along with build material can be extruded simultaneously in a FDM machine with dual extrusion hole as shown in Fig. 2. Finally, support material can be discarded either chemically or mechanically after FDM fabrication. Please cite this article as: S. Sivamani, M. Nadarajan, R. Kameshwaran et al., Ana based additive manufacturing, Materials Today: Proceedings, https://doi.org/1 An elaborate overview on the AM technologies that can be deployed for manufacturing highly complex, high value products was presented by Flaviana Calignano et al. [11]. The possibility of applying AM techniques in producing parts of an unmanned aerial vehicle were carried out by Carlo Ferro et al", " The layer corresponds to the virtual cross section of CAD model and is allowed to get fused to create the final shape. The prime advantage of this method is to manufacture complex feature or shape easily. The 3D printing method used in this study is FDM. ABS, an industrial strength plastic polymer, is utilized as a thermoplastic material. This material is fed to 3D printer in the form of filament (1.75 mm diameter) supplied by spools where it is melted in the extrusion head and feeding mechanism at around 260 degreesC (referFigure 2). Based on the earlier prepared code, the printer head lays the materials layer by layer. Thus, all the parts of CAWT are produced based on additive manufacturing concept. The parts produced by FDM method requires finishing touch. The printing surface is smoothened and irregularities if any can be removed by chemical treatment with acetone vapour, file finishing or using sand paper. The specification of additive manufacturing machine based on FDM technique available is listed in Table 3. With reference to Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002530_978-3-030-42006-2-Figure8.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002530_978-3-030-42006-2-Figure8.2-1.png", "caption": "Fig. 8.2 Local coordinate frame for a servo control unit", "texts": [ " The electronics inside the servo translates the width of the pulse into a position. A downside of this set-up is that there is no feedback from the servo, i.e. the microcontroller is unaware of the effective position of the servo. This implies a local open-loop control for the servo elements. The inverse kinematic model is based on theoretical developments from [133]. Here, an analytical method is derived for calculating the positions of the servos for a given platform position. Its schematic representation is illustrated in Fig. 8.2. As the position of the platform is known, the position of each connectionC(xc, yc, zc) is also known (as specified by the controller). The position of the servo A(xa, ya, za) and its orientation \u03b8 are determined by the motion platform construction. The connection of the servo lever and the vertical rod is indicated as B(xb, yb, zb). Let r be the length of the servo arm (AB) while BC is the shaft that connects the servo arm with the platform in the point C and has the length . The absolute distance between point A and C is denoted by 1. The coordination of point B can be characterized as xb = r sin \u03d5 cos \u03b8 + xa yb = r sin \u03d5 sin \u03b8 + ya zb = r cos\u03d5 + za (8.1) From Fig. 8.2, the following relations are found by using vector arithmetic: 21 = (xc \u2212 xa) 2 + (yc \u2212 ya) 2 + (zc \u2212 za) 2 (8.2) r2 = (xb \u2212 xa) 2 + (yb \u2212 ya) 2 + (zb \u2212 za) 2 (8.3) 2 = (xc \u2212 xb) 2 + (yc \u2212 yb) 2 + (zc \u2212 zb) 2 (8.4) By expanding (8.2) to (8.4) and inserting (8.2) and (8.3) into (8.4), we have 2 = [ 21 \u2212 ( x2a + y2a + z2a ) + 2 (xcxa + yc ya + zcza) ] +[ r2 \u2212 ( x2a + y2a + z2a ) + 2 (xbxa + yb ya + zbza) ] \u2212 2 (xcxb + yc yb + zczb) (8.5) Now expression (8.1) is inserted into (8.5), and we get 2 = 21 + r2 + [2(xa \u2212 xc)r cos \u03b8 + 2(ya \u2212 yc)r sin \u03b8 ] sin \u03d5 + [2(za \u2212 zc)r ] cos\u03d5 (8" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001696_physreve.90.052703-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001696_physreve.90.052703-Figure1-1.png", "caption": "FIG. 1. (Color online) Schematic illustration of the TaylorCouette geometry. (a) Fluid between two coaxial cylinders with radii R\u2212 and R+. The system is invariant under translations along the z-axis. (b) Planar cut perpendicular to the z-axis. The polarization vector p and the polarization angle \u03c8 with p = (cos \u03c8, sin \u03c8) are shown. From Ref. [14].", "texts": [ " The remaining coefficients \u03b6 , \u03b6\u0304 , \u03b6 \u2032, and \u03bb1 are specific for active fluids and describe, respectively, the generation of stresses and changes of the polarization field due to active processes. Note, that with our sign convention, \u03b6 \u03bc > 0 corresponds to expansive and \u03b6 \u03bc < 0 to contractive active stresses. B. An active polar fluid in the Taylor-Couette geometry Consider an active polar fluid in the Taylor-Couette geometry, that is, confined to the interstitial space between two coaxial cylinders with radii R\u2212 and R+, respectively, see Fig. 1(a). The active fluid is described by the constitutive equations given above. Since we are primarily interested in the impact of the system\u2019s activity on the dynamics, we set \u03bd1 = \u03bd\u03041 = 0. This reduces the number of terms and amounts to neglecting the elastic part of the deviatory stress tensor [30]. In a real system, notably for the actin cytoskeleton, we do not expect these terms to vanish, because for typical nematics we have |\u03bd1| > 1 [29]. Similar to the results reported in Ref. [22], however, we have not found qualitatively new states for nonvanishing values of these parameters, neither for \u03bd1 positive nor negative", " At the boundaries they add trivial constants, such that we set \u03b6\u0304 = \u03b6 \u2032 = 0 without loss of generality. We will focus on situations, where the system is invariant with respect to translations along the long axes of the cylinders. Furthermore, we confine the polarization field to the plane perpendicular to the cylinder axis. As a consequence, the field is completely determined by the orientation angle \u03c8 of the vector p with the radial direction, such that the radial and 052703-2 azimuthal components of the polarization vector are given by pr = cos(\u03c8) and p\u03b8 = sin(\u03c8), respectively, see Fig. 1(b). Now, let us decompose the molecular field h into components h\u2016 and h\u22a5 parallel and perpendicular to p. We obtain h\u22a5 = h\u03b8pr \u2212 hrp\u03b8 by taking the functional derivative of F with respect to \u03c8 , h\u22a5 = \u2212\u03b4F/\u03b4\u03c8 . The parallel component h\u2016 is a Langrange multiplier assuring |p|2 = 1. In polar coordinates r and \u03b8 , the dynamic equations then read \u2202t\u03c8 = \u2212vr\u2202r\u03c8 \u2212 v\u03b8 r (\u2202\u03b8\u03c8 + 1) + 1 2 ( \u2202rv\u03b8 \u2212 \u2202\u03b8vr r + v\u03b8 r ) + h\u22a5 \u03b3 , (7) 0 = \u03c3rr r \u2212 \u03c3\u03b8\u03b8 r \u2212 (\u2202r\u03c8)h\u22a5 \u2212 \u2202r\u03bc, (8) 0 = \u2202r (r2\u03c3\u03b8r ) r2 \u2212 \u2202r (r2h\u22a5) 2r2 + \u2202\u03b8\u03c3\u03b8\u03b8 r \u2212 (\u2202\u03b8\u03c8)h\u22a5 r \u2212 \u2202\u03b8\u03bc r , (9) 0 = \u2202rvr + vr r + \u2202\u03b8v\u03b8 r " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000919_chicc.2015.7260287-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000919_chicc.2015.7260287-Figure1-1.png", "caption": "Fig. 1: Structure of the customary inverted pendulum", "texts": [ " Section 1 introduces the structure of inverted pendulum on a cart driven by the combination of the horizontal and vertical control forces. The models of the inverted pendulum are also given. In Section 2, we give the design procedure of the PID controllers. In Section 3, the simulation results of the x-z inverted pendulum are given in different conditions. Section 4 gives the conclusions of this paper. The customary inverted pendulum on a cart driven by the horizontal control force is shown in Fig.1. In Fig.1, the control action is based on the horizontal displacements of the cart. Compared with customary inverted pendulum, the xz inverted pendulum which is mounted on a cart driven by the combination of the horizontal and vertical control forces is introduced in Fig.2. In Fig.2, the inverted pendulum is mounted on a little platform that can move up and down through a vertical force and can move with the cart in the horizontal direction. We can conclude the following remarks though comparison of two types of inverted pendulum", " The Lagrange\u2019s equations of the x-z inverted pendulum are \u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 d dt ( \u2202L \u2202x\u0307 ) \u2212 \u2202L \u2202x = Fx d dt ( \u2202L \u2202z\u0307 ) \u2212 \u2202L \u2202z = Fz d dt ( \u2202L \u2202\u03b8\u0307 ) \u2212 \u2202L \u2202\u03b8 = 0 (3) where Fx is the horizontal control force, Fz is the vertical control force, L is Lagrangian and is defined as the total kinetic energy minus the potential energy L = K \u2212 P (4) According to equations (1)-(4), the dynamic equations of the x-z inverted pendulum can be expressed as[14,15] (M + ma + mb)x\u0308 + mal cos \u03b8\u03b8\u0308 \u2212mal sin \u03b8\u03b8\u03072 = Fx (5) (ma+mb)z\u0308\u2212mal sin \u03b8\u03b8\u0308\u2212mal cos \u03b8\u03b8\u03072 = Fz\u2212(ma+mb)g (6) cos \u03b8x\u0308\u2212sin \u03b8z\u0308+l\u03b8\u0308\u2212g sin \u03b8 = 0 (7) where \u22120.5 \u2264 x \u2264 0.5 and \u22120.2 \u2264 z \u2264 0.2. It is interesting to compare the equations (5)-(7) with the equations for the inverted pendulum controlled by a pure horizontal force Fx. The equations of the customary inverted pendulum as illustrated in Fig.1 is given as following equations (M + ma)x\u0308 + mal cos \u03b8\u03b8\u0308 \u2212 mal sin \u03b8\u03b8\u03072 = Fx (8) cos \u03b8x\u0308 + l\u03b8\u0308 \u2212 g sin \u03b8 = 0 (9) In spite of the apparent similarity of formulae (5)-(7) and (8)-(9), the control design of the inverted pendulum in each case is totally different. We give the following definition x1 = x, x2 = x\u0307, x3 = z, x4 = z\u0307, x5 = \u03b8, x6 = \u03b8\u0307. Then from equations (5)-(7), we can acquired the follow- ing state equations of the x-z inverted pendulum x\u03071 = x2 (10) x\u03072 = fx2 D + d1 (11) x\u03073 = x4 (12) x\u03074 = fx4 D + d2 (13) x\u03075 = x6 (14) x\u03076 = fx6 Dl + d3 (15) where D = (M + ma + mb)mb + Mma cos2 x5, fx2 = (ma cos2 x5 + mb)Fx \u2212 ma cos x5 sin x5Fz + mambl sin x5x 2 6, fx4 = \u2212ma cos x5 sin x5Fx + (M + ma sin2 x5 + mb)Fz + (M + mb)mal cos x5x 2 6 \u2212 Mmag cos2 x5 \u2212 (M + ma + mb)mbg, fx6 = \u2212(ma + mb) cos x5Fx + (M + ma + mb) sin x5Fz + Mmal sin x5 cos x5x 2 6, d1, d2 and d3 are outer disturbances" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003430_09544070211004507-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003430_09544070211004507-Figure13-1.png", "caption": "Figure 13. Eccentricity diagram of stator: (a) non-eccentricity and (b) eccentricity.", "texts": [ " Due to the strong relevance between UMP, threephase current and PM BLDC hub motor vertical eccentricity, in this article, therefore, three-phase current and vertical eccentricity of PM BLDC hub motor are selected as input layer node. In quarter HM-EV system, the bearing between stator and rotor is Figure 9. HM-EV system model validation test bench. Figure 10. Comparison of stator vertical acceleration in frequency-domain. simplified as vertical and longitudinal spring with relatively high stiffness. Hence, the vertical eccentricity is defined as the vertical dynamic displacement between stator and rotor. The output layer node contains UMP in x-axis and y-axis direction (Figure 13). First define the sum of squares of the difference between the expected output and the actual output as the error function, and it can be formatted as follows35: er = 1 2 X i (Xm i Yi) 2 \u00f014\u00de where i is the ith neuron, Xm i is the ith actual output on the mth output layer, Yi is the ith expected value of output. BP algorithm modifies the weight coefficient according to the negative gradient direction of the error function. The weight correction formula can be derived as follows35: Dvij t+1\u00f0 \u00de= hdmj Xm 1 j +aDvij t\u00f0 \u00de \u00f015\u00de where t is incremented by 1 for each sweep through the whole set of input-output cases, Dvij is corrected weight between ith neuron and jth neuron, h is step size, dmj is balance coefficient of jth neuron on mth layer, Xm 1 j is jth actual output of m-1th layer, a is an exponential decay factor between 0 and 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.1-1.png", "caption": "FIGURE 2.1 Tire road interface.", "texts": [ " Other forces acting on the vehicle could be from external disturbances (e.g., aerodynamic forces from crosswind). However, the contact between vehicle and road is by far the dominant factor in vehicle behavior and may be the difference between safe and unsafe conditions. Therefore, emphasis is put on the influence of tire properties in general and specifically in this chapter, which describes the tire steady-state behavior. Transient and dynamic tire performance will be discussed in Chapter 3. The tire road interface is schematically shown in Figure 2.1. The tire is a complex structure, consisting of different rubber compounds, combinations of rubberized fabric, or cords of various materials (steel, textile, etc.) that act as reinforcement elements (referred to as plies) that are embedded in the rubber with a certain orientation. The outer part of the tire is cut in a specific pattern (tread pattern design), referred to as the tire profile. The tire profile serves to guide the water away from the contact area under wet road conditions, and to adapt to the road surface in order to maintain a good contact (and therefore load transfer) between tire and road" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002979_0954406220979334-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002979_0954406220979334-Figure1-1.png", "caption": "Figure 1. ARAS-Diamond robot: the novel 2RT spherical parallel robot for eye telesurgery.", "texts": [ " Toosi University of Technology, 16315-1355 Tehran, Iran. Email: taghirad@kntu.ac.ir mechanism to mechanically generate the RCM point.27 This robot is designed to perform as the slave robot in a robotic-assisted eye telesurgery system in which the surgeon operates on a master haptic console to send desired commands to this slave robot.28\u201330 However, for the sake of brevity, details of the robot design are omitted in this paper. The workspace of this robot is a hemisphere, and all joint axes intersect at the RCM point. As it is shown in Figure 1, this parallel robot provides two degrees of rotational motion and one degree of transmission motion, which is totally qualified for eye surgery. Owing to the parallel structure of the robot, higher structural stiffness, and improved position accuracy is obtained compared to that of other existing mechanisms, which makes it more appropriate for precise motions such as in eye surgeries. To provide a backdrivable transmission mechanism with high accuracy and zero backlashes, the manipulator utilizes two capstan drives for both actuators that provide rotational motions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001220_j.triboint.2014.10.021-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001220_j.triboint.2014.10.021-Figure3-1.png", "caption": "Fig. 3. Roller follower assembly instrumented with Eddy current sensor.", "texts": [ " The sensor should be small in size having high frequency response as limited space was offered by the complex shape of end pivoted roller follower for mounting of sensors. Initially, three different techniques were developed to monitor the roller rotation in the engine head. These techniques were based on Eddy current sensor, GMR sensor (Magnetometer) and Reed switch. In the first technique, Eddy current sensor Kaman U2 was used for measuring the roller rotational speed. A small hole of 4 mm was drilled at the side of the roller housing and a small nut of 5 mm was spot welded on top of the hole required to hold the Eddy current sensor in place as shown in Fig. 3. A small hole of 2 mm diameter and 0.5 mm deep was machined on the side of the roller to act as a target for the Eddy current sensor. As the roller rotated in its housing, a Transistor-Transistor Logic (TTL) signal was generated, the frequency of which was proportional to the rotational speed of the roller. One of the drawbacks in using of Eddy current sensor was that it required regular adjustment of gain and offset during tests under different lubricant inlet temperatures. In addition, the sensor can only sense the target at a Please cite this article as: Khurram M, et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001407_s11538-014-9968-0-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001407_s11538-014-9968-0-Figure10-1.png", "caption": "Fig. 10 Density plot of oxidized mediator for flow past a cylindrical anode with mixed inoculation for Reynolds number 0.1 (top). The white line indicates the surface of the biofilm. Graph showing fraction of mediator that is oxidized and fraction of biomass that is mediator utilizing along y = 0.025 cm line (bottom). The dotted line indicates the surface of the biofilm. A large flux of oxidized mediator appears on the upstream side of the biofilm, but not on the downstream side, leading to more mediator-utilizing bacteria on the leading side (Color figure online)", "texts": [ " As in the previous case, the mediator-utilizing bacteria thrive near the anode surface where the mediator has a smaller distance to travel. Near the surface of the biofilm, on the upstream side, the mediator-utilizing bacteria are able to thrive due to the steady supply of oxidized mediator from the influent. However, it has a very low density on the downstream side. This is because the mediator near the biofilm surface is already reduced before it reaches the downstream side. This is evident from the plot of the oxidized mediator in Fig. 10. We have developed a model for studying the competitive growth characteristics between anode-respiring bacteria that utilize a diffusive mediator, and bacteria that utilize an EPS matrix to transfer electrons to the anode. The purpose is to better understand the composition of biomass in a MFC reactor under various flow conditions and how that composition impacts the current generated by the reactor. This will assist in determining the optimal operating conditions for a MFC reactor to efficiently reduce the pollutants in the waste water while producing energy" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure9-1.png", "caption": "Figure 9. 1 Body: Volume = 85534 mm3", "texts": [], "surrounding_texts": [ "The numerical simulations using the ANSYS finite element software package were performed in this study for a simplified version of a disc brake system which consists of the two main components contributing to squeal the disc and the pads. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. a) Boundary conditions applied to the disk In this FE model, the disk is rigidly constrained at the bolt holes. The bolt holes are tied as rigid body to a reference point, where the rotation of the disk is allowed in the y-axis, its angular velocity is imposed and constant \u03c9= 157,89 rad/s. The support is applied to the hole of the disk and is of cylindrical type of which the degrees of freedom are shown in Matrix 1. b) Boundary conditions and loading applied to the pads Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. The boundary conditions applied to the pads are defined according to the movements authorized by the caliper. Indeed, one of the roles of the caliper is to retain the pads which have the tendency natural to follow the movement of the disk when the two structures are in contact. The caliper maintains also the plates in direction Z. Thus, the conditions imposed on the pads are: \u2022 The pad is embedded on its edges in the orthogonal plan on the contac surface, thus authorizing a rigid movement of the body in the normal direction with the contact such as one can find it in an automobile assembly of brake (Coudeyras;2009) \u2022 A fixed support in the finger pad. \u2022 A pressure P of 1 MPa applied to the piston pad. \u2022 The contact between the pads and the disk has a coefficient of friction is equal to 0.2 . Boundary conditions in embedded configurations are imposed on the models (disc-pad) as shown in Fig. 12 (a) for applying pressure on one side of the pad and Fig.12 (b) for applying pressure on both sides of the pad. Matrix 1. 1 Radial Free Axial Fixed Tangential Fixed Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited." ] }, { "image_filename": "designv11_22_0002475_s40430-020-02295-5-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002475_s40430-020-02295-5-Figure4-1.png", "caption": "Fig. 4 Motorized spindle and ceramic bearings", "texts": [ " The type of the measured bearing was H7009C, and it was assembled in a motorized spindle with the type of 150MD18Y14.5. The bearing had the same rotational speed as the rotor due to interference fit between bearing inner ring and shaft. The rated speed of the motorized spindle is 18,000\u00a0r\u00a0min\u22121. The bearing rings, balls, and cage were made of ZrO2 ceramics, Si3N4 ceramics, and PEEK resins, respectively. The structural parameters of the bearing are listed in Table\u00a01. The motorized spindle and its supporting bearings are shown in Fig.\u00a04. A water cooling system and an oil\u2013gas lubrication system were applied to the cooling and lubrication of the motorized spindle. The cooling temperature was set at a constant of 18\u00a0\u00b0C. The engine oil with 30# was used for bearing lubrication. The pressure of air was 0.28\u00a0MPa. The flow rates of lubrication oil and cooling water were set to 0.025\u00a0mL\u00a0min\u22121 Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:311 1 3 311 Page 8 of 16 and 5.0\u00a0L\u00a0min\u22121. The preload of bearings was adjusted to 350\u00a0N" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure14-1.png", "caption": "Fig. 14. Structural analysis result of pitch bearing.", "texts": [], "surrounding_texts": [ "The reliability of the structural analysis was evaluated ex- perimentally. The stress level at the same loading condition was compared analytically and experimentally at the same position of the test rig. The applied loading was an extreme load case for a pitch bearing. A strain gauge was used to measure the stress level of the test rig. Ten strain gauges of two different types were applied at key locations in the test rig, as shown in Figs. 16-18. The gauge type was chosen according to the shape of the mounted parts, and the specifications of each type are shown in Table 3. Fig. 19 shows the stress results of the different approaches. The analytical and experimental results show good agreement in terms of the tendency and magnitude of stress. Therefore, it may be concluded that the analysis method is suitable for estimating the test rig\u2019s stress level." ] }, { "image_filename": "designv11_22_0000929_romoco.2015.7219726-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000929_romoco.2015.7219726-Figure8-1.png", "caption": "Figure 8. Collision-free mobile manipulator motion.", "texts": [ " For this solution, the final time T is increased and it equals 24.03 [s], but the determined controls do not exceed the assumed constraints as can be seen in Fig. 6 and 7. In this case the way of task execution is similar to the first simulation \u2013 the mobile manipulator penetrates the safety zones of each obstacles but does not collide with them. Finally, the collision-free robot motion taking into account mechanical and control constraints and maximizing the manipulability measure is presented in Fig. 8. In this paper a method of trajectory planning in the case of the mobile manipulator has to reach a specified endeffector position in the workspace including obstacles has been presented. This approach guarantees fulfillment of mechanical constraints, obtaining continuous, limited controls and high manipulability measure in the final point of motion. Therefore, after the end of the movement the mobile manipulator is in the configuration that allows the execution of a next task without the necessity of reconfiguration" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003719_s42417-021-00346-2-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003719_s42417-021-00346-2-Figure3-1.png", "caption": "Fig. 3 Shoes and drum coupling are modelled as an interface composed several mechanical springs. By this method, the contact stiffness is calculated", "texts": [ " However, this modification would not warrant some increase in performance that justifies the mass and inertia increase [26]. Assuming unidirectional sliding and permanent contact, the Coulomb\u2019s law of friction writes: where ft is the frictional force, fn is the normal force and is the friction coefficient. According to Norton and Orthwein [26, 28] fn is expressed by the following equation: where w is the lining width. The next step is relating Eqs. (3) and (4) to obtain the frictional force along the lining: (2)pmax = a sin( max). (3)p( ) = sin( ) pmax sin( max) , (4)ft = \u2212 fn, (5)fn = pwrd , 1 3 According to Fig.\u00a0 3, the Coulomb\u2019s law can also be expressed for shear and normal stress, and , respectively: and therefore, relating Eqs. (3), 7 and 8 it can be obtained the stresses along the linings: and Ansys is the software used for this analysis and its theory to evaluate the equivalent stress is Von Misses. This calculation are accomplished by the equation: Young\u2019s modulus is a parameters that relates stress ( ) and strain ( ). Hence this lining material can be expressed by: x is the normal stress, xy is the shear stress, dL is the elongation or compression offset of the component and L is the length of it" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001325_ssd.2014.6808802-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001325_ssd.2014.6808802-Figure12-1.png", "caption": "Fig. 12. Implemented Quadrotor", "texts": [ " 1 Communicates between the ground control station and the quadrotor. Flight monitoring Bluetooth stick and/or new mission commands updates are sent via this link A computer interface application (shown in Fig. 11) was designed using visual studios and written in C# language. It used the Bluetooth protocol to check for connection and to send the appropriate commands (such as stabilize at the desired height) from the ground command station, a Laptop, to the quadrotor. Finally, the built UA V is shown in Fig. 12. Note the two controllers in the zoom-in part. The vehicle weighted 0.64 Kg, had a motor to motor distance of 330mm and used 8 inch propellers. It can carry an extra workload of 0.5 Kg. It drew a maximum of 12Amps of continuous current and l30W of continuous power. The time of flight (with a standard 3-cells, 2200mAh battery) was about 15 minutes. IV. CONCLUSION This paper presented an implementation of a modified quadrotor. An Adruino platform controller was interfaced with a GU-344 controller on a Gaui 330x Quad-Flyer in order to accept ground station commands for taking-off and stabilization at a desired height" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001387_cjme.2014.01.211-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001387_cjme.2014.01.211-Figure3-1.png", "caption": "Fig. 3. Parametric 3D dynamic model of the table", "texts": [ " The most common twin ball screw feed drive for precision positioning consists of two ball screws processed as pair, which are driven by two physically separated servomotors through flexible couplings. The ball nut is attached to the table that is constrained by four sliders to move axially on two guides, as Fig. 2 shown. It should be stated that the supporting bearings on the ball screw shaft structure are not expressed in the schematic drawing. The schematic of the mass-spring model considered here is presented in Fig. 3, in which the elements are assumed in the lumped form with 3D mechanical structure parameters. The ball screw shaft structure was endowed with axial stiffness and the guide-slides with support and lateral CHINESE JOURNAL OF MECHANICAL ENGINEERING \u00b7213\u00b7 stiffnesses. The parameter definitions of the twin ball screw feed drive system depicted in Fig. 3 are as follows. KgZ \u2014 Average support stiffness of the slider under tension and compression (N\u00b7\u03bcm\u20131), KgY\u2014Lateral stiffness of the slider (N\u00b7\u03bcm\u20131), Kl\u2014Axial stiffness of the ball screw shaft structure including the nut (N\u00b7\u03bcm\u20131), c\u2014Distance between two ball screws (mm), a\u2014Distance between two sliders on one guide (mm), uw\u2014Displacement in X direction of the gravity center of the table (m), ul1\u2014Displacement in X direction of the mount point for the nut at the outside ball screw (m), ul2\u2014Displacement in X direction of the mount point for the nut at the inside ball screw (m), uw1\u2014Displacement in X direction of the table at the mount point for the lateral nut (m), uw2\u2014Displacement in X direction of the table at the mount point for the inside nut (m), \u03b8wY\u2014Angular displacement around Y direction of the table at the gravity (rad), \u03b8wZ\u2014Angular displacement around Z direction of the table at the gravity (rad), M0\u2014Equivalent mass in axial direction of the shaft structure (kg), Mw\u2014Mass of table/work piece (kg), JwZ\u2014Yaw inertia of the table/work piece around Z direction (kg\u00b7m2), JwY\u2014Pitch inertia of the table/work piece around Y direction (kg\u00b7m2). The base coordinate system OXYZ and local coordinate system OwXwYwZw of the table are shown in Fig. 3. The original of the base coordinate system O is in the same plane with the axis lines of ball screws and at the perpendicular bisector of the four sliders. The original of local base coordinate system Ow is the gravity center of the table, whose position coordinates are (xw, yw, zw) in the base coordinate system, and the X axial is consistent with the feed direction. In the model, the ball screws offer the driving force in X direction and the guide-slides provide reaction forces in Y and Z directions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.22-1.png", "caption": "Fig. 3.22 Experimental setup: a particle moving in a tilted plane is observed from two different camera positions orientations", "texts": [ " Find out which settings for \u03b1, \u03b2, and \u03b3 bring the scanner into such an orientation that (i) \u2212\u2212\u2212\u2212\u2192 bullet1 in the image is oriented along the y-axis of the scan and (ii) the trajectories of both bullets lie in the scanned plane (Fig. 3.20). 3.7 Exercises 55 Exercise 3.2: Combining Rotation and Translation This example will provide the first step in the measurement of movements with video-based systems. In the example, the movement of a comet that moves in a plane in space is observed with a camera, as shown in Fig. 3.22. The object is a comet that circles around a planet. The data units are 107 km. The tasks for this exercise are as follows: \u2022 Read in the data from the file planet_trajectory_2D.txt, and write a program to calculate the planet velocity in the x- and y-directions. The data can be found in the scikit-kinematics package for Python users, and in the Kinematics toolbox for Matlab users, and are shown in Fig. 3.21. 56 3 Rotation Matrices \u2022 External information is provided, which tells us that the data center is 200\u00d7107 km in front of the camera, and the trajectory lies in a plane that is tiled by 30\u25e6 about the x-axis. Calculate the 3-D position of the trajectory of the particle (see Fig. 3.22). \u2022 Calculate the 3-D position pshifted of the comet in camera coordinates, observed from a satellite which has traveled 50\u00d7107 km toward the planet, and 100\u00d7107 km orthogonally to it. \u2022 Calculate the position pshiftRot (with respect to the camera), if the satellite is rotated 34\u25e6 downward. Code: C3_examples_rotmat.py: Python examples of different operations with rotation matrices, such as generating symbolic and numeric rotation matrices and calculating the corresponding rotation sequences. (p" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002701_aero47225.2020.9172614-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002701_aero47225.2020.9172614-Figure10-1.png", "caption": "Fig. 10b and Fig. 10a respectively. Otsuka and Nagatani", "texts": [ " These findings influenced the design of the Fractal Tetrahedron Assembly and the Tetracopter prototype. The rotor configuration of the Tetracopter is optimized for thrust generation, and in theory is relatively identical to a two dimensional rotorcraft in regard to efficiency. The effect of the frame of the single-propeller module on the performance of the enclosed rotor was analyzed using a computation fluid dynamics(CFD) software. The results of the analysis suggest a thrust reduction of 44% when compared to the same rotor in free flow. Viewing Fig. 10b, the rotor coincident with Plane 1 would be the only rotor to experience a 44% reduction in thrust. From the CFD results, it is shown that the lower rotors on Plane 2 of Fig. 10b would experience a thrust reduction of approximately 62% due to the flow obstruction caused by the Tetracopter\u2019s frame. A prototype of the Tetracopter was created to prove that the vehicle would fly in the given configuration. The goal of the prototype was to hover out of ground effect without any noticeable instabilities that could not be repaired through tuning the PID gains on the flight controller. The frame of the single-propeller submodule takes the shape Authorized licensed use limited to: Middlesex University", " As stated before, the frame of the single-propeller module is in the shape of a tetrahedron as seen in Fig. 12. Using the variables in Fig. 12, the length of a side of the single-propeller submodule a is equal to 244.55mm. Given a the remaining variables seen in Fig. 12, can be defined using equations (28) to (32) [5]. x = 1 3 \u221a 3a (28) d = 1 6 \u221a 3a (29) h = 1 3 \u221a 6a (30) R = 1 4 \u221a 6a (31) \u03c6 = tan\u22121( r x ) (32) The distance from the axis of the center rotor to the outer rotors, labeled as \u03bc in Fig. 10a, is approximately 130.4mm while the distance between the planes labeled in Fig. 10b as h is 168.8mm. The total mass of the vehicle is approximately 740 g. Given the degradation in thrust due to the frame of the Tetrahedral rotor-craft, at 75% motor throttle the estimated thrust to weight ratio is 1.94. The rotor configuration of a FTA allows the vehicle to tolerate motor failures. Consider a 16 rotor FTA as shown in Fig. 13a with a rotor configuration, shown from a top view, as seen in Consider the FTA, shown in Fig. 13b, in hover, meaning the only forces effecting the FTA are those created by the rotors and gravity" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002788_icra40945.2020.9197279-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002788_icra40945.2020.9197279-Figure1-1.png", "caption": "Fig. 1: A system of multiple quadcopters collaboratively transporting a cable-suspended payload system. The fixed inertial frame is denoted by {I} and the body-fixed frame of the payload is represented by {B}.", "texts": [ " The goal of PAC is to minimize the payload oscillations and that of CAC is to minimize the oscillations of cables connected with the followers to enable stable flight of the system. The performance of the proposed control architecture 978-1-7281-7395-5/20/$31.00 \u00a92020 IEEE 2253 Authorized licensed use limited to: Carleton University. Downloaded on September 22,2020 at 07:27:17 UTC from IEEE Xplore. Restrictions apply. is evaluated through experiments using two quadcopters with a cable-suspended payload and in simulation with three quadcopters. Consider a system of a cable-suspended payload being transported by multiple quadcopters collaboratively as shown in the Fig. 1. The body-fixed axes for the quadcopters are chosen as {b1i , b2i , b3i}, i = 1, 2, ...n with the third axis being perpendicular to the plane of the quadcopter and pointing upwards. The position of the payload is defined as x0 \u2208 R3 and its orientation is denoted as a rotation matrix R0 \u2208 SO(3) w.r.t frame {I}. \u03c1i \u2208 R3, i = 1, 2, ..n is the position vector from the centre of mass of the payload to the point of attachment of the cable of the ith quadcopter. qi \u2208 S2, i = 1, 2, ...n represents the attitude of the ith cable and Ri \u2208 SO(3), i = 1, 2, " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001243_s00170-015-6914-8-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001243_s00170-015-6914-8-Figure3-1.png", "caption": "Fig. 3 Forces exerted on the particle during the manipulation of cylindrical nanoparticles", "texts": [ " sin\u03b8\u2212maxsin\u03b8cos\u03b8\u2212Fxsin\u03b8cos\u03b8 \u00fe Fzcos 2\u03b8\u00fe mazcos 2\u03b8\u00feW sin\u03b3cos2\u03b8 \u00f06\u00de FY \u00bc I \u03b3\u0308 \u00feM \u03b8Y H cos\u03b3\u2212Fzsin\u03b3cos\u03b3\u2212V sin\u03b8cos\u03b3sin\u03b3\u2212mazcos\u03b3sin\u03b3 \u2212maysin2\u03b3 \u00fe Fysin 2\u03b3 \u00f07\u00de V \u00bc 1 sin\u03b8 FZ\u2212Fz\u2212W sin\u03b3\u2212maz\u00f0 \u00de \u00f08\u00de W \u00bc 1 sin\u03b3 FZ\u2212Fz\u2212V sin\u03b8\u2212maz\u00f0 \u00de \u00f09\u00de In the above relations, Fx; Fy; Fz;M \u03b8X ;M \u03b8Y ;W ;V are the components of internal forces and momentums; FX,FY,FZ are the components of dynamic forces; FT is the norm of dynamic forces; ax ,ay,az are the components of probe\u2019s acceleration; and m and I are the tip\u2019s mass and moment of inertia, respectively. In general, two groups of forces are involved in the displacement of cylindrical particles: one group is the forces exerted by the probe tip on the nanoparticle (Fu,Fv,Fw) and their resultants, which constitute the input forces, and the other group is the reaction forces created at the points of contact between probe tip and nanoparticle and also between substrate and nanoparticle, which have been designated by indexes \u201cT\u201d and \u201cS,\u201d respectively (Fig. 3). It should be mentioned that the contact forces and the friction between substrate and nanoparticle have been considered as distributed forces. By taking into account the forces involved and their geometrical relationships, the following equations for specifying the critical conditions of motion have been presented. Dynamic modes for manipulation of spherical and cylindrical nanoparticle have been widely investigated in recent researches, and in this paper, the perviousmodeling has been used [8, 10, 12, 21, 22]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000698_icpedc47771.2019.9036558-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000698_icpedc47771.2019.9036558-Figure1-1.png", "caption": "Fig. 1. Quadcopter", "texts": [ " Fast algorithms are developed for problem-solving the tasks and directing UAV for single and multi-UAV [6]. Modelling and Simulation of quadcopter using PID controller is discussed [7]. A neural network based control algorithm for precise landing is projected [8]. In this work a study on Quadcopter dynamics and its mathematical modelling is performed in order to control the quadcopter. Two PD and one PID controllers are tuned to get the desired trajectory tracking for different test cases. The quadcopter structure is shown in Fig. 1 and its coordinate frames considering the forces, angular velocities, moments and torques obtained from the four propellers are indicated in Fig. 2. A. Dynamic Model of the Quadrotor Earth frame or World frame (W) is characterized by Xw, Yw, and Zw axes where Zw points in the upward direction. The origin of the body frame (B) is connected at the center of mass of the quadcopter. ZB, perpendicular to the plane of the propellers is pointing vertically up during perfect hover where XB is coinciding with the preferred forward direction as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002274_j.mechmachtheory.2019.103748-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002274_j.mechmachtheory.2019.103748-Figure4-1.png", "caption": "Fig. 4. Split slider equivalent to: (a) Fig. 1 , (b) Fig. 2 .", "texts": [ " The discrimination between these two modes is based on the external force location. Sometimes, this distinction can be made a priori . But in most cases, the possibility of two modes can make the modelling process unnecessarily complicated. When possible, simplification should be carried out to avoid these complications. A common simplifying factor is the relatively small dimensions of the contact area. Another possibility is to dismember a single coupling into two or more couplings acting in parallel (e.g., see Fig. 4 in Section 3.2 ). The cross section geometry of the slider shown in Figs. 1 and 2 also influences the normal reaction distribution and, consequently, the normal reaction and friction force magnitudes. In general, the coefficient of friction available in the literature is for planar surfaces only. In Fig. 3 , the normal reaction distribution for an external applied force Q that passes through a cylindrical axis is illustrated. The angle of engagement of the pin with the insert is denoted by 2 \u03b2 ", " 1 and 2 , the normal reactions are represented by a single force, parallel to Q N (or Q N \u2032 and Q N \u2032\u2032 ) but arbitrarily located, in addition to a torque. If the equivalent force and torque are known, a decision regarding the operation mode can be made and, consequently, the correct normal reaction applied. However, in a network problem, these equivalent actions are dependent on the friction force which, in turn, is dependent on the respective normal reaction. To solve this dilemma, the slider can be split up into two small parallel sliders as shown in Fig. 4 . Each slider is capable of transmitting one independent force. These forces correspond to the normal reactions of Fig. 2 or are equivalent to the single normal reaction of Fig. 1 . There is no need to determine the mode of operation with the split slider. Notwithstanding, the vertical location of the friction force is dependent on the direction of the respective normal reaction (compare Q F \u2032 in Fig. 4 a and b). If the height of the slider is small, this fact can be neglected, and in other case, the location of the friction forces is dependent on the sign of the corresponding normal reaction magnitude, i.e., the sign of the force transmitted by the coupling. Consequently, Eq. (2) and (3) become non-linear. If the system is relatively small, in which only a few sign functions appear, it would be possible to find a symbolic solution, as in Section 4 . For larger problems, however, only numerical solutions are possible" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002371_j.optlastec.2020.106206-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002371_j.optlastec.2020.106206-Figure1-1.png", "caption": "Fig. 1. Diagram representation of Laser Beam Welding (LBW) setup employed.", "texts": [ " Welding trials were conducted on the sheet based on laser beam, laser speed and laser arc distance. Plasma arc clading coating was executed at 450 \u00b0C for 3 h to relieve the internal residual stress in the fusion zone. Microstructural studies were conducted at the fusion zone of as-weldments and PTAV coating to analyse the joining strength of the materials. The present study was carried out using Titanium alloy Grade 2 and Ti 6Al-4V alloy dissimilar materials with a thickness of 2 mm by laser beam welding as shown in Fig. 1. The chemical composition of dissimilar materials Titanium alloy Grade 2 and Ti 6Al-4V alloy are shown in Table 1. The microstructure studies were conducted on both dissimilar materials by using optical microscopy. The samples were prepared by wire cut electrical discharge machining process to the dimension of 150 mm \u00d7 150 mm \u00d7 2 mm. Robotics laser beam welding methods and three types of cooling process were used. Further, the primary, secondary and bottom purging gases were used in the laser beam welding process" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001445_978-3-319-06966-1_45-Figure45.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001445_978-3-319-06966-1_45-Figure45.1-1.png", "caption": "Fig. 45.1 Gearbox section", "texts": [ " The kurtogram [12, 13] is basically a representation of the calculated values of the SK as a function of the frequency f and the frequency band width Df. The kurtogram allows us to identify the frequency band where the SK is maximum, and this information can be used to design a filter which extracts the part of the signal with the highest level of impulsiveness. An envelope analysis can be performed on the filtrated signal to identify bearing faults. This technique has been already applied successfully in bearing fault detection and diagnosis [14\u201316]. A gearbox test rig was constructed to carry out the experiments. This gearbox type (see Fig. 45.1) is used as part of the transmission driveline on the actuation mechanism of secondary control surfaces in civil aircrafts. One of the bearings in this gearbox failed in an endurance test, making it an ideal candidate for this investigation where fast natural degradation of the bearing occurred. The main dimensions of the bearing and the attached bearing defect frequencies can be seen in Tables 45.1 and 45.2, respectively. The transmission was driven by an electric motor with a nominal speed of 710 rpm. An electric load motor placed at the opposite side of the test rig was used to apply the different loads used in the experiment, simulating the actual load conditions during the flight. The experiment ran continuously for 24 h a day over a duration of 36 days. Three accelerometers were mounted in the gearbox at locations identified in Fig. 45.1. The selected accelerometers (Omni Instruments model RYD81D) had an operating frequency range of 10 Hz to 10 kHz. These accelerometers were connected to signal conditioners (model Endevco 2775A) which were attached to a NI USB 6009 data acquisition device. Other than the vibration data, various parameters were monitored and stored at the same time and with the same sampling frequency: angular position of the input shaft, input and output torque, and shaft speed. The experiment started running on July 19, 2010, and the vibration measurements were taken on August 19, 2010, August 22, 2010, and finally on August 24, 2010" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000608_icems.2019.8921477-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000608_icems.2019.8921477-Figure8-1.png", "caption": "Figure 8. Bearing outer ring fault schematic", "texts": [ "64Hz 2 2 31 e r r BPI r n f f f p zf d f D The simulation experiment takes no-load as an example, and the simulation results of the speed signal are shown in Figure 6. The fault frequency fBPI and its multiplication can be observed both by the FFT and kurtosis spectrum analysis. Simulation experiments verify the feasibility of the speed method and the two signal analysis methods. V. BEARING FAULT DIAGNOSIS EXPERIMENT The bearing fault selected in this experimental platform is the outer ring fault. The experiment is carried out using the platform shown in Figure 7. The schematic diagram of the outer ring fault is shown in Figure 8. The fault motor is fixed on the base by the L-frame as the drive motor, connected to the load motor through the coupling, and the speed and current signals of the faulty motor are measured and collected by the encoder and the driver. In order to judge the fault situation more objectively, the diagnosis result of the mature vibration method is introduced as a reference. The vibration acquisition equipment is the HF30424 position vibration collector, and the vibration sensor probe is adsorbed on the outer surface of the fault motor through the data acquisition card to collect vibration signals" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001484_pssa.201532590-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001484_pssa.201532590-Figure1-1.png", "caption": "Figure 1 Schematic representation of the FT-SDCM system during screening of the Cu\u2013Ni thin film library.", "texts": [ " The chemical composition along the library was determined using energy dispersive X-ray spectroscopy (EDX, Oxford INCA). Crystallographic properties were studied using an X-ray diffraction (XRD) system (PANalytical X\u2019pert Pro) with a Cu Ka source without monochromator. The XRD measurements were done in 18 grazing incidence geometry, which rendered a spot size length of 3mm. The XpertHigh software was used for data analysis and peak identification. 2.2 Flow-type scanning droplet cell microscopy All electrochemical studies were performed using a flow-type scanning droplet cell microscope (FT-SDCM) [27]. Figure 1 shows a schematic description of the FT-SDCM setup during properties screening along the Cu\u2013Ni thin film library. The electrochemical cell body is a 3D printed polymer block, which was manufactured on a Project 6000 stereolithography machine (3D Systems, Rock Hill, USA). As observable from Fig. 1, the 3D printed electrochemical cell was designed using a full threeelectrode arrangement. An Ag/AgCl reference electrode (RE) was manufactured and assembled into a plastic screw. For this purpose, a high purity Ag wire (99.999%) was polarised to 0.1V versus standard hydrogen electrode (SHE) for 120 s and finally to 0.3V (SHE) for 600 s in a 1M HCl solution. The wire was rinsed with water and dried. To improve the mechanically stability of the deposited AgCl, the produced Ag/AgCl wire was inserted into a Teflon tube and filled with a hot mixture of 1M KCl and 4wt.% agar. The obtained RE was stored in 1M KCl electrolyte. Additional information related to the preparation of the micro-reference electrodes can be found elsewhere [28, 29]. A high purity gold wire (99.999%) acted as counter electrode (CE) and was mounted in a similar way like the reference electrode (Fig. 1). Phys. Status Solidi A 213, No. 6 (2016) 1435 www.pss-a.com 2015 WILEY-VCH Verlag GmbH & Co. KGaA, Weinheim Original Paper For a precise definition of the wetted area, the FTSDCM was used in contact mode. The cell is pressed with a defined force of 3N against the examined surface. To guarantee a proper sealing of the examined area, a sealing ring (0.5mm thick) made from a polymer rubber was used. Cyanoacrylate glue (UHUGmbH, Germany) was used to fix the sealing at the tip of the cell. Further information about the design and applications of the FT-SDCM can be found elsewhere [27]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001174_pime_proc_1961_175_037_02-Figure21-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001174_pime_proc_1961_175_037_02-Figure21-1.png", "caption": "Fig. 21. A Conneaing-rod. B Halfjournal. C Dial gauge. D Cross beams.", "texts": [ " Direction of loading . Primary cylinder pressure at Maximum oil inlet pressure to Frequency of loading cycle . Hardness of journal . . maximum load . bearing . . 33 in. 74 in. 250 to 1250 rev/min 100 h.p. at 1250 revlmin rt 100 000 Ib In vertical plane only 5000 Ib/inz 200 lb/inZ Once per revolution of journal 200 V.d.h. (Vickers dia- mond pyramid hardness) The lixture of measurement of deflection under static load The fixture for applying static loads and measuring the resulting deflections is shown in Fig. 21. It consisted of two mandrels which located the big- and small-end bearings and which were connected by cross beams to double-acting loading cylinders on either side. The big-end mandrel was made in two separate pats and these were positioned accurately in the bore by a rigid-box structure surrounding the big-end housing. A gap of $- in. was left between the two halves of the mandrel so that a scanning device could measure the profile of the bearing bore while it was loaded. Profile measurements were transferred from a probe within the assembly to a dial gauge on the outside" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003627_icedme52809.2021.00056-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003627_icedme52809.2021.00056-Figure12-1.png", "caption": "Figure 12. Schematic diagram of electromagnetic adsorption unit", "texts": [ " Hong Xiaowei and Chen Yong from the School of Mechatronics Engineering of Nanjing Forestry University also developed a new type of electromagnetic adsorption wall climbing robot in 2020, as shown in figure 11. This electromagnetic wall-climbing robot adopts sliding-rail contact technology, which can be used for large-scale steel structure wall inspection operations, and can realize the unity of robot maneuverability and adsorption by turning off the power of the electromagnetic adsorption unit. Its total weight is 6.7kg, the maximum moving speed can reach 5cm/s, and the carrying load weight is 3kg[14]. The schematic diagram of the electromagnetic adsorption unit is shown in figure 12. Two circular tracks on the left and right are installed on the inside of a single track, which are respectively connected to the positive and negative poles of the power supply. Each circular track is made of nylon material, and the upper and lower sides of the lower half of the horizontal section are pasted with copper skin. When the electromagnetic adsorption unit moves to this horizontal section with the crawler, the electromagnet is energized to generate electromagnetic adsorption force. When the electromagnetic adsorption unit departs from this horizontal section, the electromagnet loses its electromagnetic adsorption force when it is powered off" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure15.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure15.4-1.png", "caption": "Figure 15.4 Threaded inserts for plastics materials", "texts": [ " High-speed-steel dies can also be used, but care must be taken to ensure that the thread is being cut and the material is not merely being pushed aside. Threads cut directly in plastics materials will not withstand high loads and will wear out if screw fasteners are removed and replaced several times. Where high strength and reliability are required, threaded inserts are used. Inserts, which have a tapped inner hole to accept a standard threaded screw, can be self-tapping, expansion, heat or ultrasonic and can be headed or unheaded. Self-tapping inserts, Fig.\u00a015.4(a), cut their own threads, are available in brass, case-hardened steel and stainless steel from M3 to M16 and are used in the more brittle thermoplastics and in thermosets. Expansion inserts, Fig.\u00a015.4(b), are pushed into a pre-drilled or pre-moulded hole and are locked in place during assembly as the screw expands the insert. They are used in the softer plastics which are ductile enough to allow the outer knurls to bite into the plastic. They are available in brass from M3 to M8. Heat and ultrasonic inserts, Fig.\u00a015.4(c), can only be used in thermoplastics materials and are available in brass from M3 to M8. Heat inserts are used for low-volume production. The insert is placed on the end of a thermal insert tool and heated to the correct temperature, depending on the type of plastics component and pressed into a pre-drilled or pre-moulded hole. The plastic adjacent to the insert is softened and flows into the grooves to lock the insert in place. The thermal insert tool is removed and the plastic re-solidifies" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001790_s1064230715040139-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001790_s1064230715040139-Figure6-1.png", "caption": "Fig. 6. Violation of the self similarity property under the action of viscous friction proportional to velocity.", "texts": [ " In the calculation parameter c, which defines the limit value of the tractive force, was set to 20, and parameters and were selected in a way that in each case the obtained optimal trajectory would have end points at \u0410 and \u0412. The optimal trajectories are convex downward. The time of the motion along the trajectory with friction was expectedly higher than without friction. It should be noted that the self similarity property vanishes in the presence of viscous friction. For example, when the value of parameter changes in the presence of viscous friction proportional to veloc ity ( ), the terminal points of the derived optimal trajectories (Fig. 6) do not lie on the same line. It is shown in Fig. 6 that the line that passes through the terminal points of two trajectories does not pass through the terminal points of other trajectories and the origin of the coordinates. 0R \u03a8 0R 2 1\u03bc \u2261 524 JOURNAL OF COMPUTER AND SYSTEMS SCIENCES INTERNATIONAL Vol. 54 No. 4 2015 VONDRUKHOV, GOLUBEV CONCLUSIONS In the present paper, the optimal control formulas are derived for the general case with the presence of dry and viscous friction and a tractive force based on optimality conditions [5]. The expression for the optimal control does not include adjoint variables and depends only on one parameter , whose value is defined via the velocity components at the time when the terminal point of the optimal trajectory is reached" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002461_9781119662693.ch6-Figure6.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002461_9781119662693.ch6-Figure6.1-1.png", "caption": "Figure 6.1 (a) Shape memory effect cycle. (b) Morphologies of shape memory polymers. Specific morphologies can be designed to enhance the function of shape memory polymers. (a) | The formation of a zig-zag pattern enables the amplification of the movement of the polymers, relative to a rectangular film. In this shape, a relatively small change in angle (below 10\u00b0) leads to a large increase in length. (b) | Origami folding enables 2D-to-3D transformation, for example, the fabrication and reversible actuation of a crane-shaped device. Programming of an active region around a fold enables the wings to open and close in response to changes in temperature (T).", "texts": [ " SME and its similar term SCE mainly differ in the energy levels of permanent or original shape and programmed or temporary shapes. Applications of SME are promising because sometimes it is very difficult to do the same by conventional materials/technologies. In 1990, the SMM community introduced shape memory technology (SMT) [11, 12]. This led to the wide use of reshaped product design in many ways [7\u201311, 13, 14]. We can call SMM as \u201cthe material is the machine\u201d [12, 15] because it can integrate sensing and actuation functions together. Figure 6.1 shows schematic diagram shape memory process. This cycle consists of two stages. The first one is a hot stage which is of off-red background and the second is a cold stage which is of a blue background. In the hot phase, shape change occurs. At first, a fixed shaped film is subjected to heat and reshaped to a corrugated ring shape with the application of temperature (external force). This is the temporary shape that is fixed after cooling. Also, we can call this stage as a programming stage, resulted in the formation of temporary or programmed shape" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure8-1.png", "caption": "Figure 8. Faces Area = 5246.3 mm2", "texts": [ "89 / tire v\u00f9 rad s R = = (8) Total disk surface in contact with pads = 35797 mm2 (Fig.7) The external pressure between the disc and pads is calculated by the force applied to the disc, for a flat channel, the hydraulic pressure is (Oder et al.2009): [ ]1 . disc c FP Mpa A \u00ec = = (9) where cA is the surface area of the pad in contact with the disc and \u00ec is the friction coefficient. The surface area of the pad in contact with the disk in mm2 is given directly in ANSYS by selecting this surface as indicated the green color in Fig 8. In the case of a brake pad without groove, the calculation of the hydraulic pressure is obtained in the same manner. After viewing disputes meshes, by selecting the tasks applet mesh, we open the whole disc-pad on the FE model, we obtained the result of this imputation, which represented in Figures 10,11. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. The numerical simulations using the ANSYS finite element software package were performed in this study for a simplified version of a disc brake system which consists of the two main components contributing to squeal the disc and the pads" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001875_978-4-431-54595-8-Figure2.3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001875_978-4-431-54595-8-Figure2.3-1.png", "caption": "Fig. 2.3 Model of single-joint agonist-antagonist system with two PAMs: (a) initial state and (b) equilibrium state. For simplicity, it is assumed that the moment arm of a joint is constant and that the characteristics of two PAMs are the same", "texts": [ " The shift amount of the characteristic curve is associated with the pneumatic pressure P (motor command) from the controller. 3. The slope of the curve K.P/ gradually increases according to the increase of the pneumatic pressure P. These characteristics are formulated by (2.8), (2.9), and (2.10). In a biological system, a joint movement is controlled by antagonistic pairings of muscles like the triceps and biceps. To explain our novel control method using the 2 Motor Control Based on the Muscle Synergy Hypothesis 31 A-A concept, we start from considering a simple example of an agonist-antagonist system. Figure 2.3 shows the model of a single-joint arm with one pair of PAMs (agonist/antagonist). Let P1Pa be the internal pressure of PAM1, P2Pa be the internal pressure of PAM2, l1m be the length of PAM1, l2m be the length of PAM2, and rad be the joint angle. Dm is the radius of the joint. The characteristics of the two PAMs are assumed to be the same. At the initial state, set P1 D P2 D P0 and l1 D l2 D L. Consider the equilibrium-joint angle of the agonist-antagonist system with two PAMs. Let F1N be the contraction force of PAM1 and F2N be the contraction 32 H. Hirai et al. force of PAM2. The contraction forces of the two PAMs balance each other in the equilibrium state. Substituting (2.8) into the equation of F1 D F2 yields K.P1/.l1 l0.P1// D K.P2/.l2 l0.P2//: (2.14) Using the geometry constraint D D l1 L D L l2 (Fig. 2.3), D C2 L D K.P1/ K.P2/ K.P1/ C K.P2/ : (2.15) K.Pi/ (i = 1, 2) can also be rewritten as K.P i / D a1 L0 P i D kP i (2.16) by referring to (2.9) and defining P i D Pi c, where k D a1 L0 and c D b1 a1 . Using (2.15) and (2.16), is then given by D C2 L D P 1 P 2 P 1 C P 2 (2.17) D 2.C2 L/ D P 1 P 1 C P 2 1 2 : (2.18) Furthermore, if the A-A ratio R is defined as R D P 1 P 1 C P 2 ; (2.19) then the relationship between and R is expressed by a simple equation. By substituting (2.19) into (2.18), the linear relationship between and R is finally obtained as D M", " The A-A ratio R gives the relationship between the internal pressures of two PAMs (P1, P2). 2. Equation (2.20) provides the unique solution of the A-A ratio R according to the equilibrium-joint angle (or EP), although there are infinite candidates of PAM pressures that can achieve the EP. 3. The linear relationship between the EP and A-A ratio R enables us to control a joint movement in a simple way. 2 Motor Control Based on the Muscle Synergy Hypothesis 33 Next, consider the joint stiffness of the agonist-antagonist system with two PAMs (Fig. 2.3). The finite change in the joint angle from EP to C causes the restoring force driving the joint angle toward EP. Let F1 and F2 be the generated forces by PAM1 and PAM2. Two agonist and antagonist forces are given by F1 D K.P1/D (2.21) F2 D K.P2/D : (2.22) Since the radius of the joint is D, the restoring torque Nm is written as follows: D . F2 F1/D D .K.P1/ C K.P2//D 2 : (2.23) If the A-A sum is defined as S D P 1 C P 2 ; (2.24) then the joint stiffness G.D j j/Nm/rad can be written using (2.16), (2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003324_j.mechmachtheory.2021.104261-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003324_j.mechmachtheory.2021.104261-Figure17-1.png", "caption": "Fig. 17. Two tension rods mounted at the end of the manipulator.", "texts": [ " Meanwhile, locations of the additional middle links as well play an important role. Therefore, adopting additional middle link at right location is an effective method of improving the stability of the mechanism. But the number of additional middle links and stability of the mechanism are not positively correlated; more middle links would make the mechanism too heavy. Given the above, it is proposed to adopt tension components (e.g. tension spring, tension rope, tension rod, etc) at the end of the 2-dof DTPM. As shown in Fig. 17 , besides the two additional middle links, two tension rods are mounted at the end of the manipulator. They pre-tight the 2 SLiMs and suppress some idle motion accordingly. Observed from the physical testing, the stability of the entire mechanism is obviously improved. Comparing with other means for reducing idle motion, adding middle links has the advantage of easy implementation, low cost, and minimum changes to the structure. The lateral stiffness could also influences the application of the mechanism [42] " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001751_b978-1-78242-375-1.00001-0-Figure1.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001751_b978-1-78242-375-1.00001-0-Figure1.1-1.png", "caption": "Figure 1.1 Principle of operation of a chemical fuel cell (FC) based on proton transfer.", "texts": [ " FCs can, in principle, process a wide variety of fuels and oxidants, although of most interest today are common fuels, such as natural gas (and derivatives) or hydrogen, in which air is used as the oxidant. In a FC, fuel is fed continuously to the anode (negative electrode) and an oxidant (often oxygen in air) is fed continuously to the cathode (positive electrode). The electrochemical reactions take place at the electrodes to produce an electric current through an electrolyte, while driving a complementary electric current that performs work on the load. A schematic representation of a hydrogen and oxygen FC (based on an acidic electrolyte) is presented in Figure 1.1. At the anode of the FC, hydrogen gas ionizes, releasing electrons and creating H+ ion (protons), thereby releasing energy 2H2 ! 4H+ + 4e (1.1) e cathode oxygen reacts with protons and electrons taken from the anode to form At th water O2 + 4H + + 4e ! 2H2O (1.2) electrons (negative charge) flow from anode to cathode in the external circuit and The the H+ ions pass through the electrolyte. Importantly, the electrolyte should only allow proton transfer (or other ions in the case of other FC types) and not electron transfer (i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure3.13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure3.13-1.png", "caption": "Fig. 3.13 Examples of models of the data set of industrial parts. a \u201cBottle\u201d consisting of 143077 vertices. b \u201cCylinder\u201d consisting of 414722 vertices. c \u201cJoist Hanger\u201d consisting of 90 vertices. d \u201cMetal Part 1\u201d consisting of 97636 vertices. e \u201cPiston Rod\u201d consisting of 58544 vertices. f \u201cBalance Shaft\u201d consisting of 81120 vertices. g \u201cCube\u201d consisting of 728 vertices", "texts": [ " . . . . . . . . . . 22 Figure 3.8 Threat volume versus collision volume . . . . . . . . . . . . . . . . 26 Figure 3.9 The gripper used for all experiments . . . . . . . . . . . . . . . . . 26 Figure 3.10 Key Grasp Frame definition . . . . . . . . . . . . . . . . . . . . . . . 27 Figure 3.11 Example for the optimal grasp pose estimation algorithm only using the palm of the gripper . . . . . . . . . . . . 28 Figure 3.12 Models of the Mian data set . . . . . . . . . . . . . . . . . . . . . . . 29 Figure 3.13 Examples of models of the data set of industrial parts. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 30 xi Figure 3.14 Plot of the results of the RANSAM algorithm applied to isolated objects with different noise levels . . . . . . . . . . . . 31 Figure 3.15 Set of virtually scanned test scenes . . . . . . . . . . . . . . . . . . 32 Figure 3.16 Plot of the results of the RANSAM algorithm applied to the random viewpoint scenes . . . . . . . . . . . . . . . . . . . " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000834_978-3-319-16190-7_4-Figure4.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000834_978-3-319-16190-7_4-Figure4.2-1.png", "caption": "Fig. 4.2 Different stages of rotation for Euler angles. a Initial global system coordinate; b First rotation; c Second rotation; d Third rotation", "texts": [ " The sequence of rotations employed in the zxz convention starts by rotating the initial xyz coordinate system counterclockwise about the z-axis by an angle \u03c8. The resulting coordinate system is denoted by \u03be\u2019\u2019\u03b7\u2019\u2019\u03b6\u2019\u2019. In the second step, this intermediate coordinate system \u03be\u2019\u2019\u03b7\u2019\u2019\u03b6\u2019\u2019 is rotated counterclockwise about \u03be\u2019\u2019-axis by an angle \u03b8 to produce another intermediate coordinate system labeled \u03be\u2019\u03b7\u2019\u03b6\u2019. Finally, this last coordinate system is rotated counterclockwise about the \u03b6\u2019-axis by an angle \u03c3 to produce the desired \u03be\u03b7\u03b6 system of axes. The various phases of this sequence are illustrated in Fig. 4.2. The angles \u03c8, \u03b8 and \u03c3, which are the Euler angles, completely specify the orientation of the \u03be\u03b7\u03b6 coordinate system relative to the xyz frame and can, therefore, act as a set of three independent coordinates (Landau and Lifschitz 1976; Goldstein 1980). When using Euler angles, the elements of the complete rotational transformation matrix A can be obtained as the triple product of the matrices that define the planar rotations, that is, the elemental planar matrices (Nikravesh 1988) D \u00bc cw sw 0 sw cw 0 0 0 1 2 4 3 5; C \u00bc 1 0 0 0 ch sh 0 sh ch 2 4 3 5; B \u00bc cr sr 0 sr cr 0 0 0 1 2 4 3 5 \u00f04:1\u00de in which c \u2261 cos and s \u2261 sin" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003547_j.jmapro.2021.04.060-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003547_j.jmapro.2021.04.060-Figure6-1.png", "caption": "Fig. 6. Schematic illustration of SSS method: (a) subareas distribution in 3 \u00d7 3 form, (b) scanning arrangement of a typical circular ring, and (c) scanning arrangement of a rectangular ring.", "texts": [ " Connect the corresponding split points of opposite edges, and the divided quadrilateral subareas (Fig. 5(c)) are generated. Finally, overlap each subarea with the original sketch, and achieve the resultant subareas (Fig. 5(d)). Scanning sequence is significant to minimize temperature gradient and residual stress. To reduce the temperature gradient, the difference of scanning period between two adjacent subareas, called \u201cidle time\u201d, is desired to be as short as possible. A novel symmetrical scanning sequence (SSS) method is developed and demonstrated in Fig. 6. Fig. 6 (a) shows the scanning sequence and the detailed analysis is explained in Table 3. Fig. 6(b) and (c) exhibit the scanning arrangement of the typical circular ring and rectangular ring using the SSS approach. 3.2.1.2. Scanning pattern within the subarea. Contour-offset scanning mode (Fig. 7(a)) can improve the performance in a way, whereas increasing the efficiency is still a struggle due to the frequent \u201cturn onturn off\u201d of laser device [10]. This paper investigates the sequential scanning pattern (Fig. 7(b)) based on contour-offset scanning mode, with the dual scopes of improving efficiency and resultant performance" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002946_j.sna.2020.112448-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002946_j.sna.2020.112448-Figure3-1.png", "caption": "Fig. 3. Fabrication and integration steps; a) SLS based 3-D printing of the lateral sca while the magnet (3) and the mirror (4) is integrated to the lateral scanner, c) The co to the hydraulic actuator, after connection (6) of the device to the syringe pump via", "texts": [ " Finally, the nalytical modelling of the lateral scanner modes was presented lsewhere [20]. . Fabrication and Integration The device is 3D-printed using selective laser sintering (SLS) echnology with Polyamide material, which is a versatile mateial for biomedical applications. Polyamide exhibits biostability and iocompatibility, and has already been adapted for medical implant ncapsulation and insulation for moisture and corrosion prevenion [18]. SLS is an additive manufacturing technique, and involves se of lasers to sinter powdered material layer by layer [21]. As visulized in Fig. 3, the device fabrication process involves the following teps: Hydraulic actuator and lateral scanner are separately printed using SLS (via EOS Formiga 110) with Polyamide beads. The hydraulic actuator was printed with the actuator oriented vertically as shown in Fig. 3a in order to reduce the pore size formed Acetone thinned Smooth On XTC-3D is applied onto the hydraulic part to fill the pores for leakage prevention. In parallel, the magnet and the mirror are attached to the lateral scanning part. O. G\u00fcrc\u00fcog\u0306lu, I. Deniz Derman, M. Alt\u0131nsoy et al. Sensors and Actuators A 317 (2021) 112448 nner ( il (5) i a tub 5 h \u2022 The copper coil, having 0.18 mm diameter is winded (N = 200 turns) on the distal end of the hydraulic actuator, after merging of both parts (lateral scanner is screwed into the hydraulic actuator) (Fig. 3b) \u2022 5 mm diameter tubing is interconnected between the syringe pump and the device to deliver water into the hydraulic actuator. n t a 1) and the hydraulic actuator (2) b) XTC-3D is applied onto the hydraulic actuator, s winded to the hydraulic actuator and both parts are merged, d) Water is delivered ing. . Experimental Setup This section describes the details of the characterization of the ybrid hydraulic and electromagnetic actuator for 3D laser scaning. Fig. 4 illustrates the overall setup. A CMOS camera is utilized o monitor the elongation of the accordion part as a function of pplied water pressure with a syringe pump" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000310_stc.2411-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000310_stc.2411-Figure1-1.png", "caption": "FIGURE 1 General configuration of a variable speed wind turbine", "texts": [ " Consider system Equation 3 in which the uncertainties satisfy the constraint \u2016D\u03b1+1L(x,t)\u2016 < \u03b4, where \u03b4 is a known positive constant and L(x,t) represents the lumped uncertainties defined by L x; t\u00f0 \u00de \u00bc \u0394F x; t\u00f0 \u00de \u00fe \u0394G x; t\u00f0 \u00deu t\u00f0 \u00de \u00fe d x; t\u00f0 \u00de: (5) Designing the following terminal sliding mode control law: u t\u00f0 \u00de \u00bc G\u22121 n \u20acxd t\u00f0 \u00de \u2212 Fn x; t\u00f0 \u00de \u00fe D\u22121\u2212\u03b1 \u03bb p q e p=q\u00f0 \u00de\u22121 _e\u00fe ks\u00fe K sign s\u00f0 \u00de \u00fe \u03b4 ; (6) where k and K are positive constants, sign(s) is the signum function of sliding surface s, and \u03bb is a positive design constant, results in a tracking error that converges to zero in finite time. A detailed proof can be found in Dadras and Momeni.33 The general configuration of a variable speed WT is depicted in Figure 1. Here, the system's exogenous input is the wind speed vw(t) and its output is Pg(t). The pitch angle and generator torque reference inputs are, respectively, \u03b2ref(t) and Tg,ref(t). The rotor speed is represented by \u03c9r(t), whereas the generator's measured rotational speed is represented by \u03c9g(t). The generator's electric power is Pe(t). The hydraulic pitch system can be modeled as follows: _x1 \u00bc x2 _x2 \u00bc \u2212\u03c92 nx1 \u2212 2\u03c2\u03c9nx2 \u00fe \u03c92 nu y \u00bc x1; (7) where x1 x2\u00bd T \u00bc \u03b2i _\u03b2i T are the system states for pitch angle and angular velocity" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001240_icuas.2014.6842360-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001240_icuas.2014.6842360-Figure2-1.png", "caption": "Fig. 2. Pure yawing motion", "texts": [ " CD, CL and CM are the aerodynamic coefficients for drag force, lift force and pitch moment respectively The lateral dynamic generates the roll motion and, at the same time, induces a yaw motion (and vice versa), then a natural coupling exists between the rotations about the axes of roll and yaw [11]. In our case, to solve it, we have considered that there is a decoupling of yaw and roll movements [6]. Thus, each movement can be controlled independently. Generally, the effects of the engine thrust are also ignored [11]. In the Figure 2, the yaw motion is represented, which can be described with the following equations: \u03c8\u0307 = r (9) r\u0307 = N Izz (10) V\u0307y = Fy m \u2212 rVx (11) V\u0307x = Fx m + rVy (12) where \u03c8 represents the angle of yaw and r denotes the yaw rate, with respect to the centre of gravity of the airplane, N is the yawing moment and Izz represents the inertia in the z-axis. \u03b4r is the rudder deflection. Vx corresponds to the speed of the airplane in the longitudinal x-axis, Vy is the speed in the lateral y-axis, Fx describes the thrust force in the longitudinal x-axis and Fy denotes the component of the resultant lateral force on the yaxis" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure13.11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure13.11-1.png", "caption": "Figure 13.11 Circlips", "texts": [ " As the nut is screwed on, the nylon yields and forms a thread, creating high friction and resistance to loosening. deflected inwards and downwards. When the D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 13 Joining methods 13 190 bite into the shaft and cannot be removed without destroying it. This type is available up to 25 mm diameter and can be used on all types of material, including plastics. Circlips are used to lock a variety of engineering features. An external circlip, Fig.\u00a013.11(a), usually fitted in a groove in a shaft, prevents the shaft from moving in an axial direction or prevents an item fitted to the end of a shaft from coming loose, e.g. a bearing or pulley. Similarly, an internal circlip, Fig.\u00a013.11(b), can be used in a groove to prevent an item such as a bearing from coming out of a recess in a housing and will withstand high axial and shock loading. Circlips are available in sizes from 3 mm to 400 mm and larger and are manufactured from high-carbon spring steel. Lugs with holes are provided for rapid fitting and removal using circlip pliers. Smaller shafts can use a variation of the circlip, known as an E-type circlip, or retaining ring, Fig.\u00a013.11(c). These provide a large shoulder on a relatively small diameter, e.g. rotating pulleys can act against the shoulder. Wire rings or snap rings, Fig.\u00a013.11(d), can be used\u00a0as a cost-effective replacement for the traditional type of circlip in both internal and external applications such as the assembly of needle bearings and needle cages and sealing rings. where dismantling with hand tools is required for servicing and where parts are contaminated with oil. Parts can be loosened using normal hand tools without damaging the threads. Low strength is recommended for use on adjusting screws, countersink head screws and set screws on collars and pulleys and is also used on low-strength metals such as aluminium and brass which could break during dismantling" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002507_s40435-020-00640-z-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002507_s40435-020-00640-z-Figure3-1.png", "caption": "Fig. 3 Elastic displacement in L-shape mechanism", "texts": [ " (16), H lin is an identitymatrix and Dlin is a vector of zeros; therefore, the sate output vector y is equal to the state vector of x. A compliant L-shape mechanism, made of two steel rods, connected by a rigid aluminum joint is chosen as the principle of the reference mechanism (Fig. 2). The rotational motion of the first link, connected to the electric motor, can be implemented through a torque-controlled actuator. The whole mechanism can swing in the three-dimensional environment, so the gravity force affects both the rigid and elastic motion of the compliant mechanism and cannot be ignored. It can be clearly seen from Fig. 3 that the reference mechanism comprises two elements with the same length and each link is modeled with two elements. That is, the whole compliant mechanism is presented with 30 nodal elastic degrees of freedom (both linear and rotational) and one generalized coordinate (angular position q); accordingly, each finite element comprises 12 degrees of freedom. The mechanical properties of the reference mechanism are listed in Table 1. Regarding to [22], after assembling the two rods and taking into account the constraints by the kinematic coupling which needs to force zero the displacements of u1, u2, u3 and u6 in order to make the system solvable, the resulting compliant mechanism can be described by 24 nodal elastic displacements and one rigid degree-of-freedom" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002204_0954406219900219-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002204_0954406219900219-Figure6-1.png", "caption": "Figure 6. Representative macrovolume of the finger foot.", "texts": [ " Combining the geometric characteristics of the finger beam RMV with the definition of volumetric porosity and the specific surface, the parameters can be calculated by the following equations fb \u00bc Is Z b d Is \u00fe Ib\u00f0 \u00de Z b d \u00bc Is Is \u00fe Ib fb \u00bc Z 2Ib d \u00fe 2b d \u00f0 \u00de Is \u00fe Ib\u00f0 \u00deZ b d \u00bc 2Ib \u00fe 2b Is \u00fe Ib\u00f0 \u00de b 8>>< >>: \u00f014\u00de where Ib can be obtained by equation (15) while ignoring its variety along the moulded line direction Ib \u00bc D2 b D2 f 4NLst Is \u00f015\u00de Similar to the finger beam field, the RMV of the finger foot field is selected as shown in Figure 6. Its size is at a microscale, dr, in the radial direction, a finger repeat angle in the circumferential direction and a double thickness of finger laminate in the axial direction. The volumetric porosity and specific surface of the finger foot field are calculated by the following equations ff \u00bc 0\u00f0 \u00de Di N Z b dr Di N Z b dr \u00bc 1 0 ff \u00bc 2Z 2 0 r N dr\u00fe b dr 2 r N Z b dr \u00bc bN\u00fe 2 0r br 8>>>>< >>>>: \u00f016\u00de where r ranges from Di to Df, which is a narrow range, and the change in r has little influence on the specific surface of finger foot, so it is assumed that r \u00bc Di=2\u00feDf 2 2 in this paper" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure3.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure3.6-1.png", "caption": "Fig. 3.6 Visualization of the occurring problem when sheet metal parts are scanned from a single point of view. a Model of a joist hanger. b Scan of the same object. It is obvious that the scan does not have a homogeneous point density (SICK IVP Ruler E1200)", "texts": [ " . . . . . . . . . . . . 15 Figure 3.3 The assumption of a tangential contact between two oriented point pairs can be used to define a relative transformation ATB . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 Figure 3.4 Pose hypotheses generation using a \u2018birthday attack\u2019like approach. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 Figure 3.5 Scan of piston rods lying on a table (SICK LMS400) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 Figure 3.6 Visualization of the occurring problem when sheet metal parts are scanned from a single point of view . . . . . . . 21 Figure 3.7 Rotational and translational invariant features of a tripole . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22 Figure 3.8 Threat volume versus collision volume . . . . . . . . . . . . . . . . 26 Figure 3.9 The gripper used for all experiments . . . . . . . . . . . . . . . . . 26 Figure 3.10 Key Grasp Frame definition . . . . . . . . . . . . . . ", " And even worse, as the quality of the result is measured by the amount of vertices in contact, the false pose will be ranked better than any possible pose found on the real objects surface. This problem can easily be avoided by cropping the search area. But this is only one aspect of the real problem. Other drawbacks cannot be avoided as easily. For example, if objects on a table are scanned from above and these objects consist of many planar faces with 90\u25e6 angles, only edge vertices and planes are scanned and the same problem occurs (see Fig. 3.6). The problem is the same as before, namely the inhomogeneous point density. Regarding the object shown in Fig. 3.6, another problem becomes obvious. Many possible dipoles found on this object will have the same relation vector. The relation tables will not get filled properly and many false pose hypotheses are the result. This problem is known in literature as self-similarity [7]. One possibility to overcome the described issues of inhomogeneous point densities and redundant relation vectors, is to extract regions of the scan which contain most of the information and to delete all other regions. Due to unknown object poses, the alternative to fill missing points is not possible in general" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002866_0954406220971666-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002866_0954406220971666-Figure8-1.png", "caption": "Figure 8. Prototype of the lens: (a) Lens including the actuator. (b) Prototype of the proposed motor.", "texts": [ " It can be found that these two points move along the same elliptical trajectory, as the travelling wave propagates. However, their phase difference is always p in the exact same moment. That is, these two points move along the ellipse tracking in a counterclockwise direction, but their location is symmetric about the origin of the ellipse. In order to validate the feasibility and practicability of the proposed lens, the prototype with weight of 36.4 g was manufactured by SCNUAA (Nuaa Super Control Technology Co., Ltd, China), as shown in the Figure 8. The material of the lens frame is the aluminum alloy, whose inner and outer diameter is 28mm and 35mm respectively. The height of the rim is 2.8mm and it has a gap of 1mm in one side. Table 2. Materials used in the piezoelectric motor. Material Aluminum alloy PZT-8H Density (kg m 3\u00de 2810 7650 Poisson\u2019s ratio 0.33 0.3 Young\u2019s modulus (GPa) 71 120:6 53:5 51:5 53:5 120:6 51:5 51:5 51:5 104:5 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 31:3 0 0 0 31:3 0 0 0 34:6 2 6666664 3 7777775 Piezoelectric constant (C m 2) / 0 0 0 0 10:3 0 0 0 0 10:3 0 0 4:1 4:1 14 0 0 0 2 6666664 3 7777775 Besides, the flexible membranes in the lens are made of PDMS, which uses the mixing ratio of 10:1 to fabricate the membrane at 60 C" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002914_978-3-030-60330-4-Figure3.8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002914_978-3-030-60330-4-Figure3.8-1.png", "caption": "Fig. 3.8 Maps of equal frequencies for the square lattice of round particles (f = 1): a LA-mode; bTA-mode; cRO-mode. Curves 1 correspond to = 1.2; 2\u2212 = 1.5; 3\u2212 = 1.9; 4\u2212 = 2.5; 5 \u2212 = 2.65; 6 \u2212 = 2.75; 7 \u2212 = 2.85; 8 \u2212 = 2.95; 9 \u2212 = 3.1", "texts": [ "2, and all three wave modes (LA, TA, and RO) are present in the system, when 2.29 < < 3.09. It should be also noted that the maximum values of all three modes and the minimum value of the rotational mode vary insignificantly in comparison with the considered ones in the previous case. It is convenient to study isotropic and anisotropic properties of this lattice using the maps of equal frequencies. Such maps for longitudinal, transverse, and rotational phonons at different fixed values of the frequency are shown in Fig. 3.8 (for f = 1) and in Fig. 3.10 (for f = 1.5). Figure 3.8 shows that anisotropy of the longitudinal mode is most evident when > 2.7, and the longitudinal mode is almost isotropic for < 1.5, whereas the transverse mode is anisotropic even for small . The rotational mode is isotropic for 2.42 < < 2.8. However, all modes are anisotropic even in the low-frequency range in the case of the rectangular lattice with ellipse-shaped particles; see Fig. 3.10. Two modes (LA and RO) are presented at frequency = 3.1 in Fig. 3.10 in addition to Fig. 3.9, where all three modes are separately plotted in the frequency range 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001428_1350650114559617-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001428_1350650114559617-Figure5-1.png", "caption": "Figure 5. Deformation process with rollers\u2019 error.", "texts": [ " According to force equilibrium and geometric relations, the sum of the components of the load in the vertical direction of each roller is equal to the external load Q0 cos \u00feQ1 cos \u2019 \u00f0 \u00de \u00feQ2 cos 2\u2019 \u00f0 \u00de \u00fe \u00feQk cos k\u2019 \u00f0 \u00de \u00bc Fr \u00f016\u00de The sum of the components of the load in the horizontal direction of each roller should be zero Q0 sin \u00feQ1 sin \u2019 \u00f0 \u00de \u00feQ2 sin 2\u2019 \u00f0 \u00de \u00fe \u00feQk sin k\u2019 \u00f0 \u00de \u00bc 0 \u00f017\u00de Based on force equilibrium equations (16) and (17), we can judge if the offset parameters of inner ring center are correct. If it does not meet the force equilibrium, change r and , then recalculate the roller elastic deformation and load, judge the force equilibrium again. Through this iterative process, the right r and can be obtained finally and the load on each roller is calculated. Mechanics model with considering geometrical error of roller. In practice, there will be often dimension error in roller, it will influence the load distribution and contact stresses. As shown in an example in Figure 5, suppose that roller no. 0 has negative error, diameter is D0, the standard roller diameter is Dr.Then it should compare the deformation r cos \u2019 \u00f0 \u00de calculated by the deformation coordination relationship with the offset of rollers\u2019 diameters Dr D0. If r cos \u2019 \u00f0 \u00de5 \u00f0Dr D0\u00de, it indicates that the deformation between inner and outer ring is less than the offset of the rollers\u2019 diameters and roller does not contact with inner or outer raceway (as shown in Figure 5(a)), the real deformation of roller no. 0 0 \u00bc 0. If r cos \u2019 \u00f0 \u00de4 \u00f0Dr D0\u00de, it indicates that the deformation between inner and outer ring is bigger than the offset of the rollers\u2019 diameter and roller contacts with inner and outer ring (as shown in Figure 5(b)), the real deformation of roller no. 0 is 0 \u00bc r cos \u2019 \u00f0 \u00de Dr D0\u00f0 \u00de. Therefore, the real deformation of the roller k is obtained k \u00bc r cos k\u00f0 \u00de Dr Dk\u00f0 \u00de \u00f018\u00de where Dk is the diameter of roller no. k. Through numerical calculation method, the calculation process for r and is as follows: (1) Give the related parameters of bearing, calculate the contact parameters of each roller, Dr, L, E 0, i, o, Dmr, and t; (2) Give the initial value of iteration calculation r and , the increment r and ; (3) Substitute r and into equation (18), judge if each roller is loaded and calculate the elastic deformation k; (4) Based on k and equation (13), to obtain the load of each roller Qk using Matlab numerical calculation; (5) Substitute Qk into equations (16) and (17) to judge force conditions both in the horizontal and load direction, change the increments r and if force is unbalanced (choosing the force balance criteria as 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003221_s0031-8914(58)95049-3-FigureI-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003221_s0031-8914(58)95049-3-FigureI-1.png", "caption": "Fig. I.", "texts": [], "surrounding_texts": [ "THE RISING OF TOPS WITH ROUNDED PEGS 3 1 7\nThus Py = Qfl - Rt and\nPP'f l + (PQ' + QR')fl = (RR' + PT') t + PS' ,\nwhence the particular integral for fl is\nfl : (RR' + PT') t / (PQ' + QR') + PS' / (PQ' + QR').\nFor large spin this is effectively\nfl : [# -- r(h + a cos c~)]t/Cno. (8')\n3. The perturbing [orce systems. In the derived expressions for the rate of rise, #, from (8) and (8'), the terms ignored were those which differed by a factor of the order of co~no from those retained. Thus as far as the perturbation is concerned, it is natural to seek the explanation in terms of slipping, #~ and r, any other force component having to be considerably larger to produce an equivalent effect. Thus one need consider only those disturbances which give rise to these components.\nThe slipping effect has often been considered, though as to the magnitude of the effect one can say little without experimental knowledge of i Vr~\u00b0)l.\nOtherwise, one would expect on physical grounds forces at the point of contact and air effects to be the only possible sources of perturbations. On a rolling top a force parallel to k0 at P produces no resultant effect, the contributions from #~ and r cancelling. Thus, for rolling, the only effective component at P is a couple, which we may assume tentatively to be a couple /~ about the line of intersection of the rolling plane with the plane of i and z, its sense being such as to oppose rolling, giving #t = + / ' cos a.\nThe magnitude of air effects is difficult to determine theoretically. Certainly the 'r' component exists and is negative, air resistance opposing the motion. We may presume the effect to depend upon the velocity of G. For a steady rolling precession the radius, R, of the track of P is given by\nR -~ (n/co)a sin ~,\nwhence the speed of advance of G is\n(an[-- hco) sin ~.\nHence the effect would give a rate of rise decreasing with ~ and reversing its sense below the critical speed.\nAgain, air forces will give rise to a component ~,. This may be estimated by superimposing the steady precessional motion on the known solution s) for the flow in the neighbourhood of a rotating disc, taking the flow as that at the outside of the boundary layer and applying Bernoulli's Theorem to the composite flow. One determines in this way\n~, : -- 0.026 n0co sin c~ (g cm 2)", "318 D.G. PARKYN\nfor the particular top used in the experiments. This is in fact very small and would produce a fall rate of 0.00063 deg/sec, at ~ = 10 \u00b0.\nI t seems likely that air effects are entirely negligible, but it should nevertheless be possible to distinguish results due to air forces from those due to a rolling couple b y the way in which they depend upon ~. Since there is no obvious reason why F should depend upon ~, the rolling peg being spherical, the rate of rise due to roiling friction should be constant for small ~.\n4. Experimental method. If the spin, n, the inclination ~ and the radius of the peg track, R, can l~e accurately determined, it is possible to say definitely whether a top is rolling or sliding 0rod hence to decide upon the al~propriate explanation of the rise. The track may be recorded, as was done b y F o k k e r 6), b y spinning the top on paper over inverted carbon paper, but this has the disadvantage that the track produced is rather coarse.\nInstead a thin film of graphite was painted on glass sheets, using a colloidal suspension of graphite in water. On this the top left a very fine track with the graphite completely removed, which could be accurately measured with a travelling microscope. I t is not however easy to maintain a uniform thickness of graphite over a series of plates, and with a thicker film the graphite was not entirely removed. In this case it was found possible to read the tracks when illuminated from the front.\nThe spin was determined with a stroboscope, which required a voltage stabilizer in order to produce consistent results, but probably even then had an error of 4- 3% in use.\nThe angle ~ was determined b y filming the top, and then projecting the film and measuring the inclination of the top to the vertical at right and left extremes of the precession. The average of these two gave an inclination free from base error, and certainly accurate to half a-degree.\nUnfortunately it was impossible to use the stroboscope during the filming of a run because of the intensity of lighting required. The top was started", "THE RISING OF TOPS WITH ROUNDED PEGS 3 1 9\nwith a hand power drill which by means of a l~heostat gave an adjustable starting speed. The speed of a filmed track was taken as the average of two runs as nearly as possible identical with it in starting speed and inclination. The resulting spins were therefore reasonably accurate in rate of fall off with precession but involved a possibly large additive error.\nThe result of a typical run is shown in Fig. 2. The continuous line is given b y the inclination of a steadily rolling top with the observed n and R values. The points are those of the observed ~ and R values, which for a rolling top should of course lie on the continuous line. In twelve sets of observations the points did not depart from the line by more than could reasonably be accounted for by errors in n.\nThe conclusion is therefore that the top never departed far from the pure rolling configuration under the existing conditions.\nThe top used had a mass of 312 g, and C = 1700 g cm 2; A = 1085 g cm ~\"; h = 6.512 cm, the peg being a ~ in. ball bearing brazed into the shaft. The top had a critical speed of 192 rad/sec, while initial speeds ranged from 260 rad/sec to 170 rad/sec. I t was found that the initial speed had to be as low as 175 rad/sec before the top began to fall, (Fig. 3), and that even then it required a relatively large value of ~ before falling would take place..\n5. Nature o/the tracks. At high spins on a very thin graphite film the tracks showed distinct periodic oscillations (Fig. 4), in the form of clear dashes joined by dotted portions of trace. Their periodic time could be" ] }, { "image_filename": "designv11_22_0000443_ab466b-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000443_ab466b-Figure2-1.png", "caption": "Figure 2.Depositionway and sampling schematic for tensile specimens (a), and the dimension of tensile specimens (unit: mm) (b).", "texts": [ "TheLMD testwas carried out byusing an in-house 6 kWsemiconductor LMDsystem inAr atmospherewithO2 content less than 100 ppm, and the schematic of deposition systemwas shown infigure 1. The laser deposition parameterswere as follows: laser power, 2 kW; scanning speed, 11mm s\u22121; powder feeding rate, 7.8 g min\u22121; spot diameter, 4mm.Thedesigned thickness of each layerwas 0.8mm. In addition, the laser depositionprocesswas conducted in themanner of short side one-way reciprocationwith the lapping rate of 50%, the detail of deposition processwas shown infigure 2(a). In order to investigate the influence of structure anisotropy of LMDed sampleon tensile property, two sampling schemeswere selected for tensile sample preparation: sampling along theXY section (scanning direction) and sampling along theZ section (deposition direction), as shown infigure 2(a). The dimensionof tensile specimenswas shown infigure 2(b). It shouldbenoted that all the tensile testswere carried out according toBSEN ISO6892-1: 2009 [23]. The LMDedTA15 alloywas subjected to double annealing treatment, including two-stage heating, thermal insulation and air cooling, as shown infigure 3. According to the (\u03b1+\u03b2)/\u03b2 phase transition temperature of TA15 alloy, three annealing temperatures were selected in thefirst annealing process to investigate the effect of annealing temperature onmicrostructure andmechanical properties of LMDedTA15 alloy" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002328_j.mechmachtheory.2020.103873-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002328_j.mechmachtheory.2020.103873-Figure10-1.png", "caption": "Fig. 10. Contact pressure on wheel flank.", "texts": [ " Based on the simulation of hypoid gear meshing, the results are shown in case1 and case2. The bending stress for pinion and wheel are calculated by quasi-static analysis method. The wheel flank was selected to show Mises stress and contact pressure results to save the post-processing. Case 1: Face-milled hypoid gear The Mises stress for wheel in two meshing moments are demonstrated in Fig. 9 . The maximum point is in the center point of contact ellipse. The 2D and 3D diagrams of contact stress are shown in Fig. 10 . The maximum value is in the center of gear flank. The bending stress on pinion and wheel flanks are depicted in Fig. 11 (a) and (b). The positive maximum value is in the root of gear flank. The negative maximum value is in the working part of gear flank. Those stress values represent the tension-compression stress in meshing process. The root stress results of pinion and wheel are shown in Figs. 12 and 13 . The blue line is enveloped from black line in Figs. 12 ( a ) and 13 ( a ). Case 2: Face-hobbed hypoid gear The Mises stress for wheel in two meshing moments are demonstrated in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001040_j.jappmathmech.2015.04.001-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001040_j.jappmathmech.2015.04.001-Figure1-1.png", "caption": "Fig. 1.", "texts": [ "4) was accomplished by a linear transformation of the oordinates and the addition of a linear feedback to the control, after returning to the variables x, the control obtained will be constrained n an absolute value in a certain neighbourhood of zero. . Control of a multilink pendulum with two-degree-of-freedom joints We will apply the approach described above to the problem of controlling a multilink pendulum. We will consider an n-link pendulum ith two-degree-of-freedom joints, controlled by a torque applied to the first link. We will represent such a pendulum as a mechanical ystem consisting of n points masses with masses m1, . . ., mn located at the ends of absolutely rigid weightless rods with lengths l1, . . ., ln Fig. 1). The first rod is attached to a fixed support by an ideal two-degree-of-freedom joint, and the i-th rod is attached to the (i \u2212 1)-th rod, = 2, . . ., n, by the same two-degree-of-freedom joint. It is assumed that the system is in a gravitational field, which is directed opposite to he Z axis of the absolute system of coordinates OXYZ. To describe the dynamics of the pendulum we will introduce generalized coordinates in the following manner. To the i-th link of the endulum we connect the system of coordinates Oixiyizi, whose centre Oi is located at the i-th joint and whose axes are parallel to the axes \u2032 \u2032 f the absolute system of coordinates OXYZ. Along with it we introduce one more system of coordinates Oixiyi z i , whose centre is also at he i-th joint and whose Oiz \u2032 i axis lies in the same plane with the Oixi axis and the i-th link (Fig. 1). We set I.M. Anan\u2019evskii, N.V. Anokhin / Journal of Applied Mathematics and Mechanics 78 (2015) 543\u2013550 547 w r r P f p R c u p a w T t l a w here i is the angle between the Oizi and Oiz \u2032 i axes, and i is the angle between the i-th link and the Oiz \u2032 i axis. Two scalar controlling torques u1 and u2 are applied to the first link of the pendulum. They rotate the link about the Oixi and Oiy \u2032 i axes, espectively. As already noted, such a pendulum has 2n different equilibrium positions, in which some of the links are directed upward, and the emaining links are directed downward" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure12-1.png", "caption": "Figure 12. Boundary conditions and loading imposed on the disc-pads: (a) One side; (b) Two sides", "texts": [ " Thus, the conditions imposed on the pads are: \u2022 The pad is embedded on its edges in the orthogonal plan on the contac surface, thus authorizing a rigid movement of the body in the normal direction with the contact such as one can find it in an automobile assembly of brake (Coudeyras;2009) \u2022 A fixed support in the finger pad. \u2022 A pressure P of 1 MPa applied to the piston pad. \u2022 The contact between the pads and the disk has a coefficient of friction is equal to 0.2 . Boundary conditions in embedded configurations are imposed on the models (disc-pad) as shown in Fig. 12 (a) for applying pressure on one side of the pad and Fig.12 (b) for applying pressure on both sides of the pad. Matrix 1. 1 Radial Free Axial Fixed Tangential Fixed Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. A fully coupled thermomechanical model was set up to predict the temperature changes of the brake disk shape caused by axial and radial deformation. Thermal conduction and convective heat transfer where the two modes of heat transfer considered. The convection heat transfer coefficient was exposed on all surfaces of the disk, and radiative heat transfer was considered negligible(Sarip,2011)", " The dissipation of the frictional heat generated is critical for effective braking performance. Temperature changes of the brake cause axial and radial deformation; and this change in shape, in turn, affect the contact between the pads and the disk. Thus, the system should be analyzed as a fully coupled thermomechanical system. In this part, the structural and thermal analyses are coupled using ANSYS Multiphysics to determine the stress levels and global deformations of the model studied (disc label) during the braking phase under the effect of temperature. According to Fig.12 (a), the finite element model and boundary conditions embedded configurations of the model are composed of a disc and two pads. The initial temperature of the disc and pads is 20\u00b0C, and the surface convection condition is applied to all surfaces of the disc and the convection coefficient (h) of 5 W/m2\u00b0C is applied at the surface of the two pads. Indeed, the air flow stops and the produced heat, get time to transmit trough the parts. This is simulated by dropping the convection down to this value which is supposed to correspond to stagnant air" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure10-1.png", "caption": "Fig. 10. Diagram of backlash function.", "texts": [ " (46) Three completely different meshing statuses may exist: (1) non-impact status (non-separation of meshing tooth surface), (2) single-sided impact status (tooth separated, but impact is only generated on the surface of meshing tooth), and (3) double-sided impact status (tooth separated, and impact is generated on both the meshing and non-meshing tooth surfaces at the same time). In the gear dynamics model, a piecewise function is used to represent the backlash: ( ) 0 \u2032 \u2032\u2212 >\u23a7 \u23aa \u2032= \u2264\u23a8 \u23aa \u2032 \u2032+ < \u2212\u23a9 ni ni ni ni ni ni b b f b b b \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 \u03b4 (47) where \u03b4ni is the relative displacement in the direction of meshing line for each gear pair. The diagram of the backlash model is shown in Fig. 10. The gear tooth profile deviates from its theoretical position in the meshing process of gear due to the machining error of the tooth surface. Thus, the meshing status is affected, and the instantaneous transmission ratio of the gear is changed. These phenomena result in the impact between the teeth and the vibration in the transmission system. The total tangential tolerance of gear is the sum of the total cumulative pitch error and the tangential tolerance of single tooth. The reducer gear studied in this work adopts the gear grinding process to reach 5- level precision, and the corresponding tolerance can be determined by referring to the gear manual" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002488_tmag.2020.2997759-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002488_tmag.2020.2997759-Figure3-1.png", "caption": "Fig. 3: Rotational Ability of Levitator", "texts": [ " Downloaded on May 30,2020 at 01:10:09 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 levitated part 180 degrees without the need for external force application. This leads to the availability of both sides of the initially levitated part for additive manufacturing operations. This is shown in fig. 3. For the initial phase of the research, the emphasis is placed on the building of parts with simple geometries. In addition, the levitator setup is thought of as an add-on item to an existing additive manufacturing machine. The methodology of use is that the levitator setup would be placed in place of the substrate. The primary advantage of the system would be in effect when the additive manufacturing system consists of multiple material nozzles working together to accelerate the manufacturing operations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001915_cbo9781139629539.020-Figure17.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001915_cbo9781139629539.020-Figure17.1-1.png", "caption": "Figure 17.1 Structure of the cross-linked polyelectrolyte gel membrane and the reversible change of the chemical stimuli-responsive membrane in the absence (a) and presence (b) of cholesterol. (c) Faradaic impedance spectra obtained for the membrane-modified electrode in the absence (a,c) and in the presence of cholesterol (b,d) corresponding to the open and closed pores, respectively. Inset (C): The reversible changes of the electron transfer resistance, Ret, derived from the impedance spectra upon addition and removal of cholesterol. (Adapted from [49], with permission; copyright American Chemical Society, 2007).", "texts": [ " Recent advances in the development of novel signal-responsive materials [51,52] allowed their immobilization on electrode surfaces to activate and inactivate electrochemical processes reversibly by chemical or biochemical signals. In one of the examples [53], a nanostructured porous membrane prepared from a polyelectrolyte, quarternized poly(2vinylpyridine), qP2VP, cross-linked with 1,4-diiodobutane, was deposited on an ITO electrode surface and the membrane pores were reversibly closed and opened upon reacting the membrane with cholesterol and washing it out, respectively (Figure 17.1(a,b)). Cholesterol binding to the polymer chains through the formation of hydrogen bonds resulted in the swelling of the membrane and closing of the pores (Figure 17.1(b)), while washing of cholesterol from the membrane restored its initial state with the open pores, (Figure 17.1(a)). For example, swelling of the membrane caused by the uptake of cholesterol from 0.13 M solution resulted in the doubling of the dry membrane thickness (from 192 to 369 nm), while the pore diameter decreased by half (from 588 to 290 nm), resulting in a factor of 3 decrease of the membrane porosity, from 51.7% to 18.3% [53]. The reversible opening/closing of themembrane pores was visualized by atomic force microscopy (AFM) and followed by measuring the electrode electron transfer resistance for a diffusional redox probe using Faradaic impedance spectroscopy (Figure 17.1(c)). The cyclic addition and removal of cholesterol to and from the membrane resulted in a modulated interfacial resistance controlled by chemical signals (Figure 17.1(c), inset). pH-sensitive polymer brushes [51,52] tethered to electrode surfaces demonstrated pH-switchable interfacial behavior when a neutral state of the polyelectrolyte was impermeable to ionic redox species, keeping the electrode mute, while the ionized state of the polymer brush was permeable to ionic redox species of the opposite charge, allowing their access to the conducting support and activating the electrochemical process [17,54,55]. Protonation\u2013 deprotonation of the polymer brush was controlled by the pH value, which could be changed in the bulk solution [17,54,55] or varied locally at the interface by means of electrochemistry [56,57]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003523_s10846-020-01276-z-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003523_s10846-020-01276-z-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of helicopter slung load system", "texts": [ " 3) Two sufficient conditions on guaranteeing the existence of DOB state feedback controller gain and DOB output feedback controller ones are respectively proposed, and these control algorithms are established in terms of the LMIs, which can be easily tested and applied to the real cases. The rest of this work is organized as follows: a mathematical model for the helicopter with slung load and some preliminaries are presented in Section 2, the disturbance observer and model reference tracking controller are co-designed in Section 3, Section 4 gives an example to verify the validity of our methods and the work is concluded in Section 5. 2 Problem Formulation and Preliminaries In Fig. 1, a schematic diagram of helicopter slung load system is shown and the reference frames, including the inertial frame Re = {oe, xe, ye, ze}, the body fixed frame Rb = {ob, xb, yb, zb}, and the hanging frame Rh = {oh, xh, yh, zh}, are defined. As a multi-body dynamical system, the helicopter slung load system can be derived by adding the interaction forces between helicopter and load. Considering the influence of force and moment caused by the load on the helicopter, the helicopter dynamics can be written as [20] \u23a7 \u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23a9 P\u0307 = RV, mhV\u0307 = \u2212mh\u03a9 \u00d7 V + F + Fl, \u0398\u0307 = H\u03a9, Ih\u03a9\u0307 = \u2212\u03a9 \u00d7 (Ih\u03a9) + \u03c4 + \u03c4l, (1) where P = [x y z]T is the position of the helicopter in the inertial frame, and V = [u v w]T, \u0398 = [\u03c6 \u03b8 \u03c8]T, \u03a9 = [p q r]T respectively represent the translational velocity, the attitude angle, and the angular rate of helicopter in the body fixed frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001174_pime_proc_1961_175_037_02-Figure20-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001174_pime_proc_1961_175_037_02-Figure20-1.png", "caption": "Fig. 20. The fatigue test machine", "texts": [ " The machine was therefore designed to apply cyclic loading reproducing the principal characteristics only of the load diagrams of main and big-end bearings, in the form of reasonably sharp load peaks of adjustable amplitude and sharpness which could be applied to either or both sides of the bearing. The loading cycle was repeated once every revolution. The maximum peak load magnitudes were at least twice those jn the engine. The bearings were full scale. The method of applying a cyclic load to the test bearing was by making the journal eccentric and resisting the resultant motion of the bearing housing by hydraulic means. The reaction of the hydraulic forces on the bearing then provided the load. The part-sectional drawing, Fig. 19, and photograph, Fig. 20, show that the rig consisted essentially of a main shaft at the centre of which was an interchangeable bearing journal and a link-pin assembly connecting the test bearing housing to a piston. The centre sections of the shaft and link pin were eccentric with respect to their ends and the phasing of these two eccentrics determined the piston movement. The shaft and link pin rotated at the same speed but in opposite directions and were coupled together by a chain and differential gearbox, the latter being used to vary the phasing" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000315_rcar.2018.8621737-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000315_rcar.2018.8621737-Figure4-1.png", "caption": "Fig. 4. Differential drive robot model and error definition for closed loop control.e1, e2 and e3 are defined in the real robot frame, and poses of robots are defined in the world frame.", "texts": [ " Path follower \u23a1 \u23a3x\u0307y\u0307 \u03b8\u0307 \u23a4 \u23a6 = \u23a1 \u23a3cos\u03b8 0 sin\u03b8 0 0 1 \u23a4 \u23a6 \u00b7 [ v \u03c9 ] (1) \u23a1 \u23a3e1e2 e3 \u23a4 \u23a6 = \u23a1 \u23a3\u2212cos\u03b8 \u2212sin\u03b8 0 sin\u03b8 \u2212cos\u03b8 0 0 0 \u22121 \u23a4 \u23a6 \u00b7 \u23a1 \u23a3x\u2212 xr y \u2212 yr \u03b8 \u2212 \u03b8r \u23a4 \u23a6 (2) [ ue1 ue2 ] = [\u2212k1 0 0 0 \u2212sign(ur1)k2 \u2212k3 ] \u00b7 \u23a1 \u23a3e1e2 e3 \u23a4 \u23a6 (3) [ v \u03c9 ] = [ cose3 0 0 1 ] \u00b7 [ ur1 ur2 ] \u2212 [ ue1 ue2 ] (4) k1 = k3 = 2\u03be \u221a ur2(t)2 + gur1(t)2 (5) k2 = g \u00b7 |ur1| (6) After path generation, trace is passed to a path follower. We use the closed loop controller proposed by Klancar et al.[21] to make the robot move along a reference path. The robot architecture can be seen in Fig.4. For a differential drive robot, the motion equations are described by Eq.1, where v and \u03c9 are the forward and angular velocities, and \u03b8 is the forward direction of the robot in the world frame. The error between the real pose (x, y, z) and the reference pose (xr, yr, \u03b8r) in the frame of the real robot can be calculated by Eq.2. Multiplying the error by gain matrix K, we can get the feedback (ue1 , ue2) (shown in Eq.3). The final output actions(u1, u2) can be obtained from reference actions (ur1 , ur2) and (ue1 , ue2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003578_iccmc51019.2021.9418287-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003578_iccmc51019.2021.9418287-Figure3-1.png", "caption": "Figure 3. Moisture Water Sensor.", "texts": [], "surrounding_texts": [ "The power supply block consists of a step-down transformer, a bridge rectifier, capacitor, voltage regulator and rechargeable battery. By using bridge rectifier, the power supply from the main source will be step down to lower voltage range and it is rectified to get the direct or DC current. This rectified d irect current is filtered and regulated to the whole circuit operating range with a capacitor filter and voltage regulator IC. The current from the IC regulator is used to recharge the battery. This rechargeable battery acts as main power supply for supplying power to all the sections mentioned above. The power supply consists of a transformer, which will step down the input voltage 230 volts to the required operating voltage of 9 volts, thereby diodes will convert the AC current into DC and this process of conversion is called as rectification. After this process, ripples can be eliminated by passing the output through filters, which may either be capacitor or inductor. Generally , capacitor filters are most widely used rather than inductor filters. A constant or steady output voltage can be obtained by using the voltage regulators." ] }, { "image_filename": "designv11_22_0001279_s0263574714000149-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001279_s0263574714000149-Figure9-1.png", "caption": "Fig. 9. The planar four bar mechanism considered in space.", "texts": [ " Mark the permitted motions on a reference frame associated with the support; by counting their number the value of the mechanism spatiality is obtained. For illustration, we recall the example of the simple planar four bar mechanism. If considered in space it becomes a \u2018paradoxical\u2019 mechanism since it has mobility only in the very particular case of perfect parallelism between the revolute joint axes. However, KCB model reduces the problem at a straightforward open chain mobility evaluation that quickly points out the loop connectivity S = 3 (Figure 9). With S determined in this way, the four bar mechanism KCB mobility formula gives obviously a correct result irrespective if we consider it as planar or spatial mechanism. M = \u2211 fi \u2212 S = 4 \u2212 3 = 1 http://journals.cambridge.org Downloaded: 18 Mar 2015 IP address: 128.233.210.97 Theoretically, the same principle could be of course applied when the mechanism is modeled with KCB\u2019s created by opening mobile elements or joints, but in this case the redundant constraints are more difficult to identify" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure9.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure9.4-1.png", "caption": "Figure 9.4 Cam-lock spindle nose", "texts": [ "1 Centre lathe CHAPTER Turning 9 D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 9 130 9 Turning bed and contains the spindle, gears to provide a range of 12 spindle speeds, and levers for speed selection. The drive is obtained from the main motor through vee belts and pulleys and a series of gears to the spindle. The speed range is from 40 to 2500 rev/min. The spindle is supported at each end by precision taper-roller bearings and is bored through to accept bar material. The inside of the spindle nose has a Morse taper to accept centres. The outside of the spindle nose is equipped with means of locating and securing the chuck, faceplate or other workholding device. The method shown in Fig.\u00a09.4, known as a cam-lock, provides a quick, easy and safe means of securing work-holding equipment to the spindle nose. The spindle nose has a taper which accurately locates the workholding device, and on the outside diameter of the spindle nose are three cams which coincide with three holes in the face. The workholding device has three studs containing cut-outs into which the cams lock, Fig.\u00a09.4(a) and (c). The lathe bed is the foundation of the complete machine. It is made from cast iron, designed with thick sections to ensure complete rigidity and freedom from vibration. On the top surface, two sets of guideways are provided, each set consisting of an inverted vee and a flat, Fig.\u00a09.2. The arrangement shown may vary on different machines. The outer guideways guide the saddle, and the inner guideways guide the tailstock and keep it in line with the machine spindle. The guideways are hardened and accurately ground", "3 Gap bed Saddle guideways Tailstock guideways Figure 9.2 Lathe bed guideways The bed is securely bolted to a heavy-gauge steel cabinet containing electrical connections and a tool cupboard, and provides a full-length cuttingfluid and swarf tray. The complete headstock consists of a box-shaped casting rigidly clamped to the guideways of the To mount a workholding device, ensure that the locating surfaces of both parts are clean. Check that the index line on each cam lines up with the corresponding line on the spindle nose, Fig.\u00a09.4(a). Mount the workholding device on the spindle nose, ensuring that the scribed reference lines A and B on the spindle nose and the workholding device line up. These lines assist subsequent remounting. Lock each cam by turning clockwise, using the key provided. For correct locking conditions, each cam must tighten with its index line between the two vee marks on the spindle nose, Fig.\u00a09.4(b); if this does not happen, do not continue but inform your supervisor or instructor who can then carry out the necessary adjustment. Since each workholding device is adjusted to suit a particular spindle, it is not advisable to interchange spindle-mounted equipment between lathes. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 9 131 9 Turning Mounted in the dovetail slideway on the top surface of the saddle, the cross-slide moves at right angles to the centre line of the machine spindle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002913_j.optlastec.2020.106727-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002913_j.optlastec.2020.106727-Figure1-1.png", "caption": "Fig. 1. Schematic illustration of rotating CW laser treatment.", "texts": [ " The calculation results were verified with an experimental method. Furthermore, surface morphologies of laser-treated specimens under different processing parameters were adopted to investigate the surface defects resulting from laser treatment. During laser surface treatment, the CW laser beam was rotated by a prism inside laser equipment, and projected to a target surface. The laser ring was manipulated to move at a given speed and direction until covering the entire area of the target surface, as shown in Fig. 1. The laser-beam scanning speed vb in Fig. 1 refers to linear speed of rotation, and moving speed vr describes the movement of the laser ring. There were two overlapping regions during laser treatment, as shown in Fig. 1. The laser ring overlapped 50% of the ring diameter every time shift at vertical direction, the overlapping rate was calculated by dividing the overlapped width of laser ring by the ring diameter. The overlapping rate of the laser beam is represented by the ratio of overlapped width of laser beam and the beam diameter [21,22], which is changed with the processing parameters. Laser devices with power of 2 kW and 3 kW were used for surface treatment. The radius of laser beam Rb was 215 \u03bcm for both devices, and the rotation radii Rr of the 2-kW and 3-kW devices were 5 mm and 13 mm, respectively", " Optics and Laser Technology 135 (2021) 106727 TP1(z) = Tn\u2212 1(z)+ AP \u0305\u0305\u0305\u0305\u0305\u0305\u0305 2/e \u221a \u03c0\u03c1cvbRb(z + z01) (11) and TP2(z) = Tn\u2212 1(z)+ 2AP e\u03c0\u03c1cvb(z + z02) 2 (12) Finally, the peak temperature distribution considering overlapping effects can be obtained by substituting the results of Eqs. (11) and (12) into Eq. (8). The laser treatment used in this study was a rotating laser beam that traveled with the movement of the rotation center. Consequently, there were various overlapping areas including laser ring and beam overlaps, as illustrated in Fig. 1. However, the time interval of laser ring overlap was in the range of 3\u201330 s, depending on the moving speed of the laser ring. For laser beam overlap, the time interval was approximately equal to the rotation period, which was in the order of milliseconds. Considering the extremely fast cooling rate during laser irradiation (approximately 103\u2013108 K/s) [24,25], the thermal effects of repetition irradiation were dominated by a laser-beam overlapping process. Therefore, the laser-ring overlapping effects were ignored in the temperature calculation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000834_978-3-319-16190-7_4-Figure4.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000834_978-3-319-16190-7_4-Figure4.1-1.png", "caption": "Fig. 4.1 a Translation and rotation; b pure rotation", "texts": [ "eywords Euler angles Bryant angles Euler parameters The location of a rigid body i in the three-dimensional space can be defined by three translational coordinates and three rotational coordinates that describe the origin and orientation of the body-fixed coordinate system \u03bei\u03b7i\u03b6i attached to the body with respect to the global frame xyz, as it is illustrated in Fig. 4.1a. For the purpose of concentrating on the rotational coordinates of a body, let eliminate the translational coordinates by allowing the global and local coordinates systems to coincide at the origin, as Fig. 4.1b shows (Nikravesh 1988). The orientation of a rigid body i can be specified by a transformation matrix, the elements of which may be expressed in terms of suitable sets of coordinates, such as Euler angles, Bryant angles or Euler parameters. Since the motion of the body is continuous, the transformation matrix must be a continuous function of time. It is known that the nine direction cosines present in the rotational transformation matrix Ai expressed in Eq. (3.7) define the orientation of the \u03bei\u03b7i\u03b6i axes" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000549_0731684419888588-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000549_0731684419888588-Figure1-1.png", "caption": "Figure 1. Different type of chiral unit cell.15\u201317,32", "texts": [ " To present the deformation mode and elastic properties, the finite element analysis (FEA) model is developed under quasi-static compression, and the effects of geometric parameters on effective Young\u2019s modulus, relative twist angle, and effective Poisson\u2019s ratio are discussed. Finally, quasi-static compressive experiments are carried out to verify the FEA model and show the compression-twist effect. The common type of chiral 2D cells includes three, four, and six chiral ligaments, which is shown in Figure 1. In these cells, on the one hand, Poisson\u2019s ratio of tri-chiral cell is always positive, Poisson\u2019s ratio of anti-tri-chiral cell with short ligament is negative, and Poisson\u2019s ratio of anti-tri-chiral cell with long ligament is positive.16 On the other hand, tetra-chiral and hex-chiral have stable Poisson\u2019s ratio, which is close to 1.15 For effective Young\u2019s modulus, its value could increase with the increase of cell number,17 and for out-of-plane elastic properties, chiral structures with more ligaments have better transverse shear modulus" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003394_s10846-021-01352-y-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003394_s10846-021-01352-y-Figure7-1.png", "caption": "Fig. 7 MAV mass moment of the inertia test rig", "texts": [ " Applying the designed thrust test rig in the wind tunnel, the thrust for two RPMs is plotted versus speed from 5 m/s to 15 m/s, as shown in Fig. 6. As expected, the thrust decreases as the wind speed increases and increases with the propeller speed. An accurate mass moment of inertia is required for the flight dynamics simulation. This requires measuring the mass moment of inertia of the MAV. Fundamentally, this requires suspending the MAV from two points equidistant from its center of gravity, as shown in Fig. 7. The MAV is rotated by a small angle about its center of gravity and released. The oscillation period is measured, and the moment of inertia can be calculated. Further details regarding this technique can be found in [41]. A wooden stand is designed for this purpose. The measured mass moment of inertia is Iyy = 0.0007 Kg.m2. To validate the output from the nonlinear simulation, a test rig is built to measure theMAV angle change with respect to time in the wind tunnel. This technique has been used by Balakrishna and Niranjana [42] and Pattinson et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001239_0954410014537240-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001239_0954410014537240-Figure15-1.png", "caption": "Figure 15. Flight experiment stand. (a) Framework and (b) physical graph.", "texts": [ " The simulation results show that the proposed controller has smaller tracking error and better performance compared with the existing method. Several flight experiments have been carried out to investigate the control performance of the proposed scheme on the real helicopter. To facilitate experiment implementation and to guarantee the safety of the at University of Sydney on March 13, 2015pig.sagepub.comDownloaded from helicopter, we construct a five degrees of freedom (DOFs) flight experiment stand in this paper,29 as shown in Figure 15. The stand is a mechanical construction able to hold a helicopter, allowing basic movements while protecting it from damaging and crashing. The roll, pitch, and yaw axes provide the whole DOFs of rotation, and the main and elevation axes provide two DOFs of translation. The helicopter is fixed on the stand when flying, where the motion of it will be affected. In this paper, the constraint of the stand is treated as one part of the total disturbances, which is similar with the parametric uncertainties and un-modeled dynamics can be estimated by the extended state estimator (equation (29))" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001483_0954406215612814-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001483_0954406215612814-Figure8-1.png", "caption": "Figure 8. Generation of the base circle B as envelope of the normal lines N .", "texts": [ " Base circle of involute circular gears The base circle of circular gears is the evolute of each involute tooth profile and can be determined as the envelope of the family of all straight lines that are normal to the rack-cutter profile and pass through the instant center of rotation P0. In fact, when the normal line N to both profiles at the contact point P passes through P0, they are conjugate, since the relative sliding velocity is tangent to them at point P. From previous work,6 extended here, this family of normal lines N is generated by the pure-rolling motion of the auxiliary centrode \" that coincides with the pitch line of the rack-cutter, whose motion parameter is the angle , measured positive clockwise, as shown in Figure 8. The position vector N of a generic point Q of the normal line N in the moving frame O2 X2 Y2, whose X2-axis coincides with the auxiliary centrode \" or pitch line of the rack-cutter, can be expressed as N 2 \u00bc rp \u00fe r cos , r sin T \u00f04\u00de where rp, r, and are the pitch radius of the pitch circle , the oriented segment from P0 to Q, the clockwise angle that gives the position of P0 with respect to the Y1-axis, and the pressure angle of the rack-cutter profile , respectively. The rotation matrix R from the moving frame O2 X2 Y2 to the fixed frame O1 X1 Y1 attached to the gear of pitch circle , can be expressed, in frame-1, as R\u00bd 1\u00bc cos sin sin cos \u00f05\u00de The translation vector s of O2 with respect to O1 in frame-1 is given by s\u00bd 1\u00bc u v \u00bc rp sin rp cos rp cos \u00fe rp sin \u00f06\u00de which can be also expressed in the form s\u00bd 1\u00bc R\u00bd 1 s\u00bd 2; s\u00bd 2\u00bc rp rp \u00f07\u00de at UNIV CALIFORNIA SAN DIEGO on December 14, 2015pic", " The foregoing procedure proves that the base circles of circular gears and, in general, those of noncircular gears, as shown by Figliolini and Angeles,6 can be obtained as the envelopes of the family of lines N normal to the rack-cutter profile , but this approach is quite complex. On the contrary, the approach proposed here makes use of Aronhold\u2019s first theorem and the return circle, which eases the formulation and its application. In fact, the same result of equation (13) can be obtained by simply tracing the return circle R of diameter equal to the pitch radius rp of the gear to generate the base circle, as shown in Figure 8. Thus, the intersection of N with R gives the point at UNIV CALIFORNIA SAN DIEGO on December 14, 2015pic.sagepub.comDownloaded from T of the base circle B. Upon repeating the procedure for different positions of the pitch line of the rackcutter, one obtains the base circle. Base curves of involute non-circular gears The approach proposed above for circular gears can be extended to the case of non-circular gears by giving their base curves as sets of points coming from the intersection of the return circle R with the line N normal to the rack-cutter profile , which is defined by the pressure angle " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure7-1.png", "caption": "Figure 7. Contact area selected ANSYS 1 Face: Ac = 35797 mm2", "texts": [ " It is supposed that 60% of the braking forces are supported by the front brakes (both rotors), that is to say 30% for a single disc (Mackin et al.2000). The force of rotor for a typical vehicle is calculated using the vehicle data contained in Table 3, resulting in: Working forces to the brake disc: ( ) [ ] 2 0 20 0 130% . 2 1047,36 12. . 2 disc rotor stop stop tire stop Mv F N R vv t t R t = = \u2212 (7) The rotational speed of the disk is calculated as follows: 0 157.89 / tire v\u00f9 rad s R = = (8) Total disk surface in contact with pads = 35797 mm2 (Fig.7) The external pressure between the disc and pads is calculated by the force applied to the disc, for a flat channel, the hydraulic pressure is (Oder et al.2009): [ ]1 . disc c FP Mpa A \u00ec = = (9) where cA is the surface area of the pad in contact with the disc and \u00ec is the friction coefficient. The surface area of the pad in contact with the disk in mm2 is given directly in ANSYS by selecting this surface as indicated the green color in Fig 8. In the case of a brake pad without groove, the calculation of the hydraulic pressure is obtained in the same manner" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003053_iros45743.2020.9341428-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003053_iros45743.2020.9341428-Figure4-1.png", "caption": "Fig. 4: LEGO structure described by a product situation", "texts": [ " A planning situation s \u2286 A, on the one hand, describes a consistent set of attributes (i. e., cons(s)) that are constructed at a given stage of assembly. The set of all situations is called S. On the other hand, a set of product attributes is given as product situation p \u2286 A Product if it describes a certain product state during planning. For a planning situation s, its corresponding product situation is defined as s Product . It is S Product = {s Product | s \u2208 S} the set of all product situations. An example of a LEGO product state described by a product situation is given in Fig. 4. Four valid product attributes (details see below) establish the structure, each connecting two bricks. The overall aim of assembly planning is to find robotic tasks which result in a planning situation that represents the completely assembled product according to a goal product situation. In this final planning situation, the specific constellations of the actuators do not matter to the reached goal. Instead, only product related parts of the planning situation count for the assembly process. A planning situation s \u2208 S fulfils a product situation p \u2208 S Product , if and only if s and p contain the same product attributes; which is s p exactly if s Product = p", " For the planning on macro level, a special form of task is introduced: The set of domain tasks t is defined as T Product so that ExplEffect+(t) \u2286 A Product and ExplEffect\u2212(t) \u2286 A Product . That is, domain tasks do only construct or remove product attributes. A task t \u2208 T \\ T Product that also influences actuator attributes is called automation task. Besides the attributes explicitly defined by a task, other attributes may also be affected by a task\u2019s execution. Assume a task explicitly constructing attribute 2 in Fig. 4. Having attributes 3 and 4 already assembled before, also attribute 1 is constructed by the task. In a product situation where none of the attributes exists before, however, only attribute 2 is constructed by the same task. Which attributes actually are modified depends on the situation the task is executed in. Given a task t \u2208 T and a situation s \u2208 S, Effect+(t, s) and Effect\u2212(t, s) describe disjoint sets of attributes that are actually constructed or removed by t in s, with: ExplEffect+(t) \u2286 Effect+(t, s) and ExplEffect\u2212(t) \u2286 Effect\u2212(t, s)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000698_icpedc47771.2019.9036558-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000698_icpedc47771.2019.9036558-Figure2-1.png", "caption": "Fig. 2. Free body diagram of quadcopter", "texts": [ " Modelling and Simulation of quadcopter using PID controller is discussed [7]. A neural network based control algorithm for precise landing is projected [8]. In this work a study on Quadcopter dynamics and its mathematical modelling is performed in order to control the quadcopter. Two PD and one PID controllers are tuned to get the desired trajectory tracking for different test cases. The quadcopter structure is shown in Fig. 1 and its coordinate frames considering the forces, angular velocities, moments and torques obtained from the four propellers are indicated in Fig. 2. A. Dynamic Model of the Quadrotor Earth frame or World frame (W) is characterized by Xw, Yw, and Zw axes where Zw points in the upward direction. The origin of the body frame (B) is connected at the center of mass of the quadcopter. ZB, perpendicular to the plane of the propellers is pointing vertically up during perfect hover where XB is coinciding with the preferred forward direction as shown in Fig. 2. By rotating the two propeller pairs on the two arms with same speed yet the other direction, the rotational torque can be dropped. Since quadcopter has just four autonomously controllable actuators but six degrees of freedom, the system is an under actuated system. These four control variables must be 116978-1-7281-2414-8/19/$31.00 c\u00a92019 IEEE Authorized licensed use limited to: Fondren Library Rice University. Downloaded on May 18,2020 at 11:21:07 UTC from IEEE Xplore. Restrictions apply. selected such that the controlling of the quadcopter will become less complicated" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001235_j.acme.2014.11.003-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001235_j.acme.2014.11.003-Figure1-1.png", "caption": "Fig. 1 \u2013 The dynamic model of the gearing [4\u20136].", "texts": [ " The first approach has been published in 1990 [1], in which the excitation forces in the gear pump has been defined and the vibration model of the gearing has been described. The current papers regarding the dynamical behavior of gear pumps describe the flow and pressure distribution inside the gear pumps, excitations of gears due to the pressure and dynamical forces in the bearings are calculated [3]. These models include also the influence of the pressure ripples inside the pump, but as stated, this influence can be neglected at higher pressure. nd dynamic loads in external gear pumps, Archives of Civil and 03 an & Partner Sp. z o.o. All rights reserved. In Fig. 1 the dynamic model of the gear pump shown in [4\u20136] has been described. This model includes the influence of the gearing on the reaction forces in the bearings. The torque on the gears changes periodically depending on the position of the teeth contact point, which is sealing the delivery area from the suction area. In Fig. 2, two characteristic positions of this point have been shown at which the change of the torque as well as the radial forces on the gears occurs. In the range 0 < \u2019 \u00f0p=z\u00de \u00f0e 1\u00de the trapped volume is connected with the pressure chamber" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000485_etfa.2019.8869079-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000485_etfa.2019.8869079-Figure9-1.png", "caption": "Fig. 9. FreeCAD-based custom workbench.", "texts": [ " Once the AutomationML file has been created, it is possible to visualize the layers, tracks and the process parameters in a CAD environment using our proposed FreeCAD-based custom workbench. The AutomationML file is the only input for this workbench and it allows the user to modify the process parameters intuitively. Such a CAD environment was created in FreeCAD software [24]. This software has the possibility to create a workbench with custom features programming in Python and C++ language by means of Qt libraries. The workbench provides functionalities to select and edit a given layer or track, Fig. 9. This is especially useful for complex part geometries where we could anticipate eventual local problems on the manufacturing of the part and then strategically modify the parameters on the desired track or layer. As the parameters are changed, the AutomationML file is also updated with the new updated parameters, which will later be used to generate the path planning or to simulate the AM process. Regarding to the offline stage, the AutomationML file contains all the information to build the part and therefore to simulate the building strategy in a robot simulation environment or in a structural simulation environment" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003324_j.mechmachtheory.2021.104261-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003324_j.mechmachtheory.2021.104261-Figure2-1.png", "caption": "Fig. 2. Typical configurations in the three workspace components (one middle link on the center-right).", "texts": [ " In the following, relationship among the workspace components of the D-SLiM is discussed. With given \u03b21 \u2208 [1 \u25e6, 179 \u25e6] , \u03b22 \u2208 [1 \u25e6, 179 \u25e6] and \u03b21 + \u03b22 < \u03c0, for the 2-dof DTPM in Fig. 1 , three workspace components are obtained by solving \u03b8 with (13) in the domain [ 0 \u25e6, 9 0 \u25e6] . The initial values are given as \u03b81 = \u03b82 = ( \u03c0 \u2212 \u03b21 \u2212 \u03b22 ) / 2 . They correspond to the configurations that the moving platform is always parallel to the ground, i.e. \u03b3 = 0 . This workspace component is denoted by Q 1 in Fig. 2 (a). Thereafter, a perturbation is imposed on the initial value. According to the positive and negative values of \u03b3 , the solutions are divided into 2 categories. When \u03b3 < 0 , the moving platform is skewed to the right. The workspace component is denoted by Q 2 , as shown in Fig. 2 (b). The other workspace component, in which the moving platform is skewed to the left ( \u03b3 > 0 ), is denoted by Q 3 , Fig. 2 (c). Fig. 3 (a) shows the three different workspace components. Variations of \u03b3 , \u03b81 and \u03b82 with respect to \u03b21 and \u03b22 are illustrated in Fig. 3 (b)\u20133 (d). As can be seen from these figures, the 3 workspace components are close to each other at singular configurations. The multifurcation phenomenon occurs after the mechanism passing through its singular positions [38,39] . When the middle link is shifted from the center-right to the center-left of the manipulator, due to the identity of the two D-SLiMs, the three workspace components of the manipulator can be derived from the previous results", " The resultant angle deviation is defined as \u03b8 \u2223\u2223 i, j = \u221a ( \u03b81 \u2223\u2223 i, j )2 + ( \u03b82 \u2223\u2223 i, j )2 , (16) where \u03b81 \u2223\u2223 i, j represents the deviation of \u03b81 in Q i and Q j with respect to the same \u03b21 and \u03b22 , \u03b82 \u2223\u2223 i, j describes the deviation of \u03b82 in different workspace components. Intuitively, the smaller the resultant angle deviation between two configurations is, the more likely the mechanism sways between them and causes idle motion. Considering the deployable manipulator with one middle link on center-right in Fig. 2 , the resultant angle deviations \u03b8 | 1 , 2 and \u03b8 | 1 , 3 are plotted in Fig. 6 . Fig. 6 shows that when \u03b21 \u2192 0 and \u03b22 \u2192 0 , both \u03b8 | 1 , 2 and \u03b8 | 1 , 3 are small. In this region, the margin between Q 1 , Q 2 and Q 3 is narrow. It would be easy for the mechanism to \u201cslide\u201d from one configuration to the other under the influence of joint clearances, and to cause the idle motion of the moving platform. However, in most region, the deviations between two configurations are not small, especially for \u03b8 | 1 , 3 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002730_cnm.3400-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002730_cnm.3400-Figure5-1.png", "caption": "Fig. 5. The model of ADAMS simulation", "texts": [ ", 5) corresponding to a set of input variables \u03d5(0, y, z, \u03b1, \u03b2, \u03b3), where li indicates the length of each strut. The parameters of the mechanism are shown in Table. 1. This article is protected by copyright. All rights reserved. ADAMS software is a widely used and authoritative kinematics and dynamics simulation software that can simulate and analyse the kinematics and dynamics performance of complex mechanical systems. The 3D mapping software Solidworks is used to model the bone-fixator system and import it into ADAMS software, as shown in Fig. 5. By giving the length of each strut of the external fixator, ADAMS software is used to verify the correction trajectory under the following three different adjustment strategies. The corresponding constraints and motion are applied to the external fixator. (1) Applying fixed constraints: the proximal ring of the external fixator is fixed with the ground, and other configured parts are fixed with corresponding parts according to the actual situation. (2) Applying rotational constraints: there are two spherical joints, six revolute joints and four universal joints in the mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002387_acs.langmuir.0c00398-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002387_acs.langmuir.0c00398-Figure1-1.png", "caption": "Figure 1. Schematic diagram of the measurements explanation of interaction forces between switchable surface-active silica particles.", "texts": [ "3 \u00b5m, which is totally safety size to prevent the overlap of laser beam for the 5.0 \u00b5m particles in this work. Moreover, a pre-installed device that can make the wavelengths of the two laser beams are inconsistent has been well designed to eliminate the interference of the overlap of the laser beams in the optical tweezers instrument even though the laser beam overlap occurs under special circumstances. The typical process of trapping and measuring the interaction forces between a couple of silica particle is presented in Figure 1. More details can refer to our previous studies41-43. Page 7 of 25 ACS Paragon Plus Environment Langmuir 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 8 According to previous studies44-45, the hydrophobic force and the steric barrier can be negligible under the experimental conditions of two identical silica particles coated with small molecule surfactant in this work. Moreover, the hydrodynamic force cannot be considered due to the quite low velocity (0", " Obviously, the EDL repulsive force between silica particles in the 1:1 molecular ratio CTAB/SDS solution began to emerge in a closer surface-to Page 16 of 25 ACS Paragon Plus Environment Langmuir 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60 17 surface distance compared with that in CTAB alone solution, which clearly illustrated that the Debye length did not switch as the switching of surface active to surface-inactive in the molecular ratio of 1:1 CTAB/SDS solution. It can be further confirmed that the cationic surfactant CTAB molecules detached from the negative charged silica particles surface to form complex or ion pair with the added anionic surfactant SDS (at 1:1 molecular ratio) molecules, as illustrated in Figure 1. Herein, a novel approach was developed to measure the interaction forces between a couple of switchable surface-active colloid particles in situ by using dual-laser optical tweezers. All the measured interaction forces can be well-fitted with the theoretical model originated from the DLVO theory, which indicated that the experimental data determined by optical tweezers was quite precise, and the theoretical model was appropriate under this work\u2019s condition. A strong repulsion increased dramatically as two silica particles approached to the minimum surface distance to engender the interaction forces, which was mainly contributed to the EDL repulsive force" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001122_cjme.2014.0918.153-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001122_cjme.2014.0918.153-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of the test apparatus", "texts": [ " The aim of this paper is to experimentally study seal face wear of a S-DGS operating under high pressure only due to the pressure deformation, XU Jing, et al: Experiment on Wear Behavior of High Pressure Gas Seal Faces \u00b72\u00b7 in which face structure is considered as one of the major factors affecting seal deformation and the measurement of optical images and profiles of sealing surface is proposed so as to provide evidence for the theoretical study. A test rig for examining seal face wear of a S-DGS operating under high pressures was designed as shown in Fig. 1. The stator of the tested S-DGS was fixed into a metallic holder. The rotor the tested S-DGS was fixed into another metallic holder on the axle sleeve of the shaft, which was driven by a motor at a certain rotation speed. An eddy current displacement sensor was fixed together with the stator and a stainless steel sheet was fixed into the rotor so as to measure the average gas film thickness. The seal clearance was got by measuring the displacement change between the eddy current displacement sensor and the stainless steel sheet during testing" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002701_aero47225.2020.9172614-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002701_aero47225.2020.9172614-Figure7-1.png", "caption": "Figure 7: Loads applied on the truss members for the hovering in the single point top attachment scenario. Payload is equal to 30 kg, or about 2.53 times the assembly mass. Blue colors indicate compressive forces while red colors indicate tensile forces.", "texts": [], "surrounding_texts": [ "for every i \u2208 {1 . . . 4n+1}, these recursive relations can be written\nMa n+1 = 14 \u2297Ma n\nM b n+1 = 2nr [[p1]\u00d7|[p2]\u00d7|[p3]\u00d7|[p4]\u00d7] (I4 \u2297Ma n)\n+ 14 \u2297M b n\nM c n+1 = 14 \u2297M c n\nMd n+1 = 14 \u2297Md n\n(5)\nwhere 1p = [11 . . . 1] is a matrix of size 1\u00d7 p for p \u2265 0 and [p]\u00d7 is the matrix corresponding to the left cross-product by p.\nFinally, by defining\nQ = \u23a1 \u23a2\u23a3 [p1]\u00d7Ma 0 0 0 0 0 [p2]\u00d7Ma\n0 0 0 0 0 [p3]\u00d7Ma\n0 0 0 0 0 [p4]\u00d7Ma\n0\n\u23a4 \u23a5\u23a6 , (6)\nwe can write\nMa n = 14n \u2297Ma 0\nM b n = 14n \u2297M b 0\n+ rQ n\u2211 k=0 2k (14n\u2212k \u2297 I4 \u2297 14k \u2297 I4)\nM c n = 14n \u2297M c 0\nMd n = 14n \u2297Md 0 .\n(7)\nThe matrices Ma 0 , M b 0 , M c 0 , Md\n0 are given by a linearization of the dynamics of the elementary module. The derivation of the dyanmics is made in the case of the Tetracopter in section 3. One can notice that the norms ofMa\nn ,M c n, andMd n\nare proportional to 4n while the dominant term in M b n grows proportionally to 8n. Therefore, by looking at the moments equation for the n-assembly in which quadratic terms are omitted:\nJn\u03a9\u0307 \u2248MT +MC ,\nif the inertia is replaced using ((2)), and if the moments are replaced by their linearization around an equilibrium point with only the dominant terms as n grows kept,\n2 9 16nmr2 (4I3 \u2212M) \u03a9\u0307\n\u2248 rQ n\u2211 k=0 2k (14n\u2212k \u2297 I4 \u2297 14k \u2297 I4)\u0394u (8)\nand one can see the left side of the equation grows proportionally to 16n while the right side grows with 8n. As n grows bigger, the dynamical system linearized equations can be seen as an expanded version of the reduced dynamical system and the system is expected to roughly behave similarly, with a characteristic time constant multiplied by 2n.\nInternal forces of the fractal assembly\nThe analysis of the internal forces of the fractal assembly structure can be made by representing the fractal assembly of 4n tetrahedra as a truss of mn members and jn joints. Since\neach tetrahedron is made of 6 members, mn = 6\u00d7 4n while the number of joints, found by induction on the size of the assembly, is given by jn = 2(4n + 1).\nTherefore we have\nmn + 6\u2212 3jn = 0\nand the truss is statically determinate internally. The direct stiffness method can be applied to evaluate the inertnal loads applied on each member in different scenarios. A similar analysis is done on the two dimensional Sierpinski gasket truss in [6]. The authors of the mentioned publication show that the members internal forces of the truss do not differ much from the forces present in the complete triangle tessellation truss of same dimensions.\nIn the three dimensional cases, different simulations are performed to evaluate the rigidity of the assembly and the potential structural failures that could occur. The truss members parameters come from the type of tube used for the construction of the Tetracopter frame. The loads on the 2- assembly are computed with the direct stiffness method in three different scenarios:\n\u2022 At rest lying on the plane only supporting its own weight; \u2022 Hovering a payload attached to the top elementary module; \u2022 Hovering a payload attached to three different attachment points at each of the bottom corners of the assembly.\nIn the hovering situations, the payload weight is gradually changed from to 0 kg to 30 kg, which represents about 2.53 the weight of the 2-assembly. Fig. 5 shows the maximum compressive and tensile loads applied to the members of the truss. During all scenarios, the displacements of the nodes remain negligible (under 10\u00d7 10\u22126 m), such that the structure deformation is minimal.\nFigures 6 to 8 show the trusses and the loads applied on its members in the three different scenarios. For the two payload lifting scenarios, the payload is set to 30 kg. We can observe that the top members are subject to higher loads than the other members. However, they still remain significantly under their critical buckling load, equal to 659N when computed with a length factor of 2.\nThe Tetracopter is represented as a rigid body and its dynamics are derived with the Newton-Euler equations. The derivation of the nonlinear dynamic equations does not differ much from the derivations found in other papers that model flat quadcopters [8, 11, 3, 15]. However, the particular placement of the rotors implies a different expression of the torque induced by the differential thrust in (22).\nWe consider an inertial reference frame and a body-fixed frame as shown on Fig. 9. The position and orientation of the body-fixed frame in the inertial frame is given by the translation vector \u03be and the Euler angles \u03b7 defined by (9) and (10).\n\u03be = [ x y z ] (9)\nAuthorized licensed use limited to: Middlesex University. Downloaded on September 02,2020 at 09:13:47 UTC from IEEE Xplore. Restrictions apply.", "(b) Three-point bottom attachment scenario\nThe rotation matrix from the inertial frame to the bodyfixed frame is given by (11), where the sines and cosines are\nabbreviated.\nR = [ c\u03c8c\u03b8 c\u03c8s\u03b8s\u03c6 \u2212 s\u03c8c\u03c6 c\u03c8s\u03b8c\u03c6 + s\u03c8s\u03c6 s\u03c8c\u03b8 s\u03c8s\u03b8s\u03c6 + c\u03c8c\u03c6 s\u03c8s\u03b8c\u03c6 \u2212 c\u03c8s\u03c6 \u2212s\u03b8 c\u03b8s\u03c6 c\u03b8c\u03c6 ] (11)\nThe body-fixed frame has linear velocity V B = [u, v, w] and angular velocity \u03a9 = [p, q, r].\nThe transformation matrix S is used to obtain the angular velocities in the inertial frame from the angular velocities in\nAuthorized licensed use limited to: Middlesex University. Downloaded on September 02,2020 at 09:13:47 UTC from IEEE Xplore. Restrictions apply.", "the body-fixed frame.\n\u03a9 = S\u03b7\u0307 = [ 1 0 \u2212s\u03b8 0 c\u03c6 c\u03b8s\u03c6 0 \u2212s\u03c6 c\u03b8c\u03c6 ]\u23a1 \u23a3\u03c6\u0307\u03b8\u0307 \u03c8\u0307 \u23a4 \u23a6 (12)\nThe forces acting on the Tetracopter in the body-fixed frame are gravity, the thrust of the four rotors, and the air drag [3]. Their sum is equal to the centrifugal force and the derivative of the linear momentum in the body-fixed frame:\nmV\u0307 B +\u03a9\u00d7 (mV B) = mRTG+ T \u2212 \u23a1 \u23a3kxu2kyv 2\nkzw 2\n\u23a4 \u23a6 . (13)\nThe scalars kx, ky , and kz are drag coefficients.\nThe thrust produced by the rotors can be written\nT = kT\n[ 0 0\n\u03c92 1 + \u03c92 2 + \u03c92 3 + \u03c92 4\n] (14)\nwhere kT is a coefficient that depends on the ambient air density and the rotor blades\u2019 characteristics and \u03c9j is the angular velocity of rotor j.\nThe angular velocity in the body-fixed frame \u03a9 is given by the Euler\u2019s equation for a rigid body\nM = Iq\u03a9\u0307+\u03a9\u00d7 (Iq\u03a9) (15)\nwhere Iq is the inertia tensor of the Tetracopter and M the applied torques.\nThe applied torques include:\n\u2022 Mj for j \u2208 1, 2, 3, 4, the counteracting torques induced on the four stators by spinning the rotor; \u2022 MT , the torques induced by the differential thrust of the rotors; \u2022 MD , the drag torque induced by the rotation of the rotorcraft around \u03a9.\nFor a rotor j, Mj can be determined with Euler\u2019s equation applied to the rotor in the body-fixed frame\n\u2212Mj +Md j +M f j = Ir\u03c9\u0307j + Ir\u03a9\u00d7 \u03c9j (16)\nwhere we use:\n\u2022 Md j , the drag torque induced by the rotation of the rotor; \u2022 M f j , the friction torque of the rotor with with the stators; \u2022 Ir, the inertia of a rotor around its rotation axis.\nSince rotors 1 and 3 rotate counter-clockwise and rotors 2 and 4 rotate clockwise,\n\u03c9j = (\u22121)j+1\u03c9je B z , (17)\nMd j = (\u22121)jkD\u03c92 je B z , (18)\nand\nM f j = (\u22121)jkF\u03c9jeBz (19)\nwhere kD and kF are respectively a drag and a friction constant [11, 3].\nCombining equations (16) to (19) gives\nMj = (\u22121)j [\nIr\u03c9jq \u2212Ir\u03c9jp\nIr\u03c9\u0307j + kD\u03c9 2 j + kF\u03c9j\n] . (20)\nThe torque induced by the drag on the rotorcraft is given by (21), where kp, kq , and kr are drag coefficients [3].\nMD = \u2212 \u23a1 \u23a3kpp2kqq 2\nkrr 2\n\u23a4 \u23a6 . (21)\n(22) gives the torque induced by the differential thrust of the rotors. a is the length of the side of tetrahedral frame and the assumption is made that the z-axis intersects the base of the tetrahedron formed by the rotorcraft in its center.\nMT = akT\n[ (\u03c92\n3 \u2212 \u03c92 1)/4\n((\u03c92 1 + \u03c92 3)/2\u2212 \u03c92 2)/(2\n\u221a 3)\n0\n] . (22)\nBy replacing M by \u22114 j=1 Mj + MT + MD in (15), we obtain\n\u03a9\u0307 = I\u22121 q ( 4\u2211 j=1 (\u22121)j [\nIr\u03c9jq \u2212Ir\u03c9jp\nIr\u03c9\u0307j + kD\u03c9 2 j + kF\u03c9j\n]\n\u2212 \u23a1 \u23a3kpp2kqq 2\nkrr 2\n\u23a4 \u23a6\n+ akT\n[ (\u03c92\n3 \u2212 \u03c92 2)/4\n((\u03c92 1 + \u03c92 3)/2\u2212 \u03c92 2)/(2\n\u221a 3)\n0\n]\n\u2212\u03a9\u00d7 (Iq\u03a9) ) .\n(23)\nAuthorized licensed use limited to: Middlesex University. Downloaded on September 02,2020 at 09:13:47 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_22_0000439_j.mechmachtheory.2019.103612-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000439_j.mechmachtheory.2019.103612-Figure6-1.png", "caption": "Fig. 6. CGs and representative distances of the arm and forearm.", "texts": [ " (8) \u2013(10) must be solved for each instant of time in order to know the position of the limb segments over time. The numerical differentiation of the position coordinates ( y, y s , \u03b8 a , \u03b8b ) can be computed to obtain the velocities ( \u0307 y, \u02d9 ys , \u02d9 \u03b8a , \u02d9 \u03b8b ) and the accelerations ( \u0308y , y\u0308 s , \u03b8\u0308a , \u03b8\u0308b ). Next, for the inverse dynamical analysis, the acceleration of the center of gravity (CG) of each limb segment is necessary. The velocities of the CGs of the arm, \u02d9 rga , and forearm, \u02d9 rg f , can be written as (see Fig. 6 ): \u02d9 rga = \u02d9 ys + \u02d9 \u03b8a \u00d7 r Aga (11) \u02d9 rg f = \u02d9 rB + \u02d9 \u03b8b \u00d7 r Bg f (12) \u02d9 ys = \u02d9 ys \u02c6 j \u02d9 \u03b8a = \u02d9 \u03b8a \u0302 k \u02d9 \u03b8b = \u02d9 \u03b8b \u0302 k r Aga = L ap cos \u03b8a \u0302 i + L ap sin \u03b8a \u02c6 j r Bg f = L f p cos \u03b8b \u0302 i + L f p sin \u03b8b \u02c6 j \u02d9 rB = \u02d9 ys + \u02d9 \u03b8a \u00d7 r AB and, r AB = L a cos \u03b8a \u0302 i + L a sin \u03b8a \u02c6 j Finally, the accelerations of the CGs of the arm, r\u0308 ga , and forearm, r\u0308 g f , can be obtained from: r\u0308 ga = y\u0308 s + \u03b8\u0308a \u00d7 r Aga + \u02d9 \u03b8a \u00d7 ( \u0307 \u03b8a \u00d7 r Aga ) (13) r\u0308 g f = r\u0308 B + \u03b8\u0308b \u00d7 r Bg f + \u02d9 \u03b8b \u00d7 ( \u0307 \u03b8b \u00d7 r Bg f ) (14) The bench press exercise involves several muscles groups to generate the movement of the upper limbs, mainly the pectoralis major, the anterior deltoid brachial and the triceps brachii [60] . In a simplified mathematical model of the muscle action, the sum of the contribution of all individual muscles in a joint is reduced to a net joint moment. In Fig. 6 , M 2 and M 3 are the net joint moments at the shoulder and elbow, respectively; F 2 is the vertical force at shoulder, which corresponds to the force usually measured with force plates located under the user\u2019s back [21,55] ; and F T is the vertical force acting on the user\u2019s wrist. In the model the net joint moment at the wrist has been neglected. It must be noted that the mechanism is overactuated by the moments at the shoulder and elbow. This mean that there is an infinite number of combinations of M 2 and M 3 that allows lifting the load along the same trajectory", " While the force value calculated adding the upper-limb inertia or the force directly measured from force plates are used in the power equation P = F \u00b7 v , it must be noted that those forces are not applied to the barbell and therefore the power calculated as the product of that force and the velocity of the barbell is less meaningful in a physical sense. In this work, based on the mathematical model of the bench press exercise, it is proposed to calculate the mechanical power as the product of the net joint moment by the angular velocity of the corresponding upper-limb segment, plus the product of the vertical force in the shoulder by its vertical velocity. The equation of power can be written as, see Fig. 6 : P = | M 2 \u02d9 \u03b8a | + | M 3 \u02d9 \u03b8 f | + | F 2 \u0307 ys | (22) where the absolute value | \u00b7 | is taken to account for the muscle power. During the raise of the load, the mechanical power is positive. However, it must be noted that during the descent of the load, the net mechanical power of the mechanism is negative, that is, the muscles need to be contracted to work as a brake, developing power. Finally, due to the assumption of M 3 \u2248 0, the power is evaluated as P = | M 2 \u02d9 \u03b8a | + | F 2 \u0307 ys | . The mathematical model of the bench press exercise described in the previous section is used for a comparative between the isoinertial resistance and the constant-force bench press proposed in this work" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure6.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure6.12-1.png", "caption": "Figure 6.12 Parallax error", "texts": [ " Neck type \u2013 for outside diameter measurement such as behind a shrouded recess. d. Tube thickness type \u2013 for tube or pipe wall thickness measurement. 1. Before use, clean the measuring surfaces by gripping a piece of clean paper (copier paper is ideal) between the jaws and slowly pull it out. Ensure the instrument reads zero before taking a measurement. Always ensure digital models are zeroed. 3. Look straight at the vernier graduations when making a reading. If viewed from an angle, an error of reading can be made due to parallax effect (Fig. 6.12). D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 6 Measuring equipment 6 87 These are also available as an easy-to-read dial depth gauge (Fig. 6.16) and a digital model with an LCD readout (Fig. 6.17) operating in the same way as the caliper models. be taken to allow for the thickness of the jaw, depending on whether the attachment is clamped on top of or under the jaw. The thickness of the jaw is marked on each instrument. Height gauges are available in a range of capacities reading from zero up to 1000 mm", "29 Measuring tube wall thickness (a) Blade micrometer (b) Tube micrometer D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 6 Measuring equipment 6 92 Micrometers with an adjustable measuring force are available for applications requiring a constant low measuring force such as measuring thin wire, paper, plastics or rubber parts which avoids distortion of the workpiece. These are available in a range from 0\u201310 mm to 20\u201330 mm. 1. Before use, clean the measuring surfaces by gripping a piece of clean paper (normal copier paper is ideal) between the anvil and the spindle and slowly pull it out. Ensure the instrument reads zero before taking a measurement. Always ensure digital models are zeroed. 2. Look straight at the index line when making a reading. If viewed from an angle, an error of reading can be made due to parallax effect (Fig. 6.12). 3. Ensure the surface being measured is clean and free of swarf and grit before taking a measurement. 4. Do not rotate the thimble using excessive force. 5. Use the ratchet device, if available, to ensure a consistent measuring force. 6. Be careful not to drop or bang the micrometer such as would cause damage to the instrument. 7. Always store the instrument in a clean place when not in use, preferably in a case or box. 8. Never leave the measuring faces clamped together when not in use. Always leave a gap\u00a0between the measuring faces (say 1 mm)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003676_s00170-021-07413-8-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003676_s00170-021-07413-8-Figure10-1.png", "caption": "Fig. 10 The assembled finite element model of the die components", "texts": [ " To simulate the HDDRP with inward flowing liquid process, the commercial FE code Abaqus/Explicit solver was utilized which is generally implemented for solving dynamic, non-linear large-scale deformation processes, including quasi-static sheet metal forming problems [35]. By considering the symmetry of the process and the material mechanical properties, a 3D elastoplastic model of the process consisting of only one quarter of the sheet and the die set was designed to save CPU time, as illustrated in Fig. 10. According to the physical nature, the tooling components (punch, die, and blank holder) were considered as rigid with 4- node elements, R3D4. The blank was considered as a deformable body and was meshed as a 3D deformable axisymmetric mesh with C3D8R 8-node elements. Five elements along the thickness were considered, and the blank model contained 9070 elements. The friction between the contact surfaces was modeled using Coulomb law [36], and the coefficient of friction was selected from previous works [8], which have the closest FEM results to the present experimental tests" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001323_ihmsc.2014.92-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001323_ihmsc.2014.92-Figure1-1.png", "caption": "Figure 1. Robots are uniformly distributed in the arena.", "texts": [ " Balancing means that when the number of robots cannot meet the total of threshold values, the system stabilizes in a state that each aggregation is filled the same percentage of its threshold value of the aggregation[13].Meanwhile, the robot needs to move among the aggregations when the number of aggregations is changed, which is called redistribution. Redistribution is the basis of keeping the equality of each aggregation. II. SELF-ORGANIZED PROBABILISTIC MODEL In this paper, based on the cockroach aggregation behavior[8], each robot is divided into two states: searching and aggregating. When all the robots are in searching state, they are distributed uniformly in the arena (Fig.1). In each state, the robot has unique behavior feature which achieves state transition through the decision-making according to comparing probability with a random number (Fig.2). A. Individual behavior Searching state is very important for the robot to achieve searching target in the workspace. In this paper, the correlated random walk inspired by random behavior of cockroaches is regarded as the motion of robots in searching state[8]. Due to the limited ability of processor chips in robots, the algorithm of random walk is required as simple as possible", " The direction of robot\u2019s velocity will rotate\u03b8 counterclockwise on the basis of the current direction of movement. Then the robot starts a new linear motion and records the length path. When the robots\u2019 distance are very close, they can obtain the relative location by infrared range and bearing system[14], then the robots determine the potential field to change the direction of velocity to prevent collision[5]. If the robot detects the boundary of the arena by infrared distance sensors, it returns to the central zone with a random angle drawn uniformly between 17 and 78 degree [11] (Fig.1). In swarm robotics, each robot has a unique code as its own identity (ID). The robot moves with limited ability of interaction including infrared sensing for local positioning and wireless communication. When the robot in searching state has the requirement to form the aggregation, it will change its own state into aggregating. The ID of the aggregation is the same as the ID of the first member of the aggregation. Sensing range of the robot in the same aggregation forms a whole, in which the robots coordinate move and transfer message with each other" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003249_tte.2021.3049466-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003249_tte.2021.3049466-Figure2-1.png", "caption": "Fig. 2. Optimization of the rotor.", "texts": [ " Thus, measures from three aspects are taken in this part. First, the DFCW instead of SFCW is adopted. In contrast with SFCW, DFCW provides a lower air-gap harmonic under the same sinusoidal phase excitation [24]. The physical and thermal isolation of DFCW can be realized by inserting insulating materials between adjacent windings in the same slots and limiting short-circuit current, respectively. Second, the optimization is conducted on both sides of the rotor frame to reduce cogging torque. As shown in Fig. 2, \u201cO\u201d is the center of shaft and PMs. Ro is the actual radius of outer rotor arc whose center has been reduced by ho from \u201cO.\u201d Ri is the actual radius of inner rotor arc whose center has been expanded by hi from \u201cO.\u201d \u201cOO \u201d and \u201cOi \u201d are center of outer and inner rotor arcs, respectively. Moreover, several magnetic bridges with enough width, about 0.8 mm, are designed to reduce the flux leakage between adjacent PMs. The cogging torque and torque versus hi and ho are shown in Fig. 3. In Fig. 3(a), it can be drawn that the cogging torque decreases first and increases slightly than with the augment of hi and ho" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure7-1.png", "caption": "Fig. 7. ROMAX model of transmission system.", "texts": [ " The radial and axial stiffness of the bearing rollers and bearing enclosure can be expressed as 2 2 ( ) ( ) cos ( ) ( ) ( ) ( ) sin ( ) ( ) \u00d7\u23a7 =\u23aa +\u23aa \u23a8 \u00d7\u23aa =\u23aa +\u23a9 c ij oil ij rij ij c ij oil ij c ij oil ij aij ij c ij oil ij K K K K K K K K K K \u03b1 \u03b1 2 2 ( ) ( ) cos ( ) ( ) ( ) ( ) sin ( ) ( ) \u00d7\u23a7 =\u23aa +\u23aa \u23a8 \u00d7\u23aa =\u23aa +\u23a9 c oj oil oj roj oj c oj oil oj c oj oil oj aoj oj c oj oil oj K K K K K K K K K K \u03b1 \u03b1 (40) where \u03b1ij and \u03b1oj are the contact angle between the bearing rollers and enclosure. By combining Eqs. (39) and (40), the equation of the radial and axial stiffness of rolling bearing can be expressed as 2 1 1 cos = = \u00d7\u23a7 =\u23aa +\u23aa \u23a8 \u00d7\u23aa =\u23aa +\u23a9 \u2211 \u2211 n rij roj r j j rij roj n aij aoj a j aij aoj K K K K K K K K K K \u03c8 (41) where n is the number of bearing rollers and \u03c8j is the angle of the j th bearing rollers. The transmission system model, as shown in Fig. 7, is established in ROMAX software and comprehensively considers the force of the shaft, the force of the bearing, and bearing clearance. The force and deformation of each part of the system are obtained through the static performance analysis. Finally, the bearing stiffness in X, Y, and Z directions of each rolling bearing, as shown in Table 3, is obtained by combining it with the calculation formula of rolling bearing stiffness. The stiffness and damping of all bearings on each shaft shall be equivalent to the coordinate origin of each moving member in the dynamic model, and the bearing damping shall be calculated by linear damping, that is, = +C M K\u03b1 \u03b2 , (42) where M is the weight of bearing, K is the stiffness of bearing, and \u03b1 and \u03b2 are the damping proportional coefficients" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000308_s11837-019-03674-7-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000308_s11837-019-03674-7-Figure1-1.png", "caption": "Fig. 1. Description of SLS process.", "texts": [ " However, studies of the oxidation of the Ti-6Al-4V alloy manufactured by SLS have not been reported to date. The aim of the present work is to investigate the oxidation behavior of the Ti-6Al-4V alloy manufactured by SLS and to reveal the growth mechanism of the oxide scales. SLS is a rapid manufacturing process which generates complex 3D objects by solidifying successive layers of powder materials. Solidification is obtained by fusing or sintering selected areas of the successive powder layers with high thermal energy supplied by a laser beam, as shown in Fig. 1. The Ti6Al-4V alloy was manufactured by SLS, and its chemical composition is listed in Table I. The specimens were cut from a turbine blade into coupons of 15 mm 9 15 mm 9 2 mm. The sampling position is shown in Fig. 2. A hole with a diameter of 2 mm was drilled near the edge of each specimen to JOM https://doi.org/10.1007/s11837-019-03674-7 2019 The Minerals, Metals & Materials Society facilitate suspension within a heating furnace with an accuracy of 1 C. All the specimens were polished by 240#, 600#, 1000#, and 2000# SiC papers followed by 5 min of ultrasonic cleaning and degreasing in acetone" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001792_s11015-016-0183-0-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001792_s11015-016-0183-0-Figure2-1.png", "caption": "Fig. 2. Diagram of the operation of a KURABO Mazerustar kk250 mixer.", "texts": [ " The molten metal used in our investigation was copper alloy of grade M1 and the atomizing gas was air. The re- sulting micropowder was subjected to treatment in a planetary mixer. In the treatment, two 10-g samples were placed in the mixer and 0.5 g of commercial-grade carbon nanopowder was added. In addition to powders of copper and carbon, the mixer contained zirconium oxide (ZrO2) in the form of spherical particles with a diameter of 1.2\u20131.5 mm and a combined weight of 10 g. The mixing operation was performed in a KURABO Mazerustar kk250 mixer (Fig. 2), with the platform rotating about the main axis at a speed of 1350 rpm while the container rotated about the local axis at a speed of 1230 rpm. The mixing operation lasted 5 min for sample No. 1 and 15 min for sample No. 2. The contents of the mixer were periodically cooled every minute (the copper powder was heated to 60\u00b0C by friction.). The intensive mixing of the powder resulted coating of the surface of the copper particles by carbon, with the remain- ing carbon staying in the free state. Sample No" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001737_j.ast.2016.03.028-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001737_j.ast.2016.03.028-Figure1-1.png", "caption": "Fig. 1. Three-dimensional pursuit-evasion geometry.", "texts": [ " For verifying the 3D tracking performance of this proposed method, several tough testing scenarios are arranged based on three random targets\u2019 maneuvers and geometry relationships between missiles and targets. From the simulation results, this proposed method delivers really promising contributions for missile terminal guidance law designs. Three-dimensional (3D) pursuit geometry of a tracking missile and a maneuverable target in terminal phase is generally described in the spherical coordinates as in Fig. 1. Fig. 1 is a pursuit-evasion geometry between the missile and target where the relative position vector along the line of sight is expressed by r . The 3D relative velocity can be obtained by differentiating r as \u0307r = r\u0307 er + r\u03b8\u0307 cos\u03c6 e\u03b8 + r\u03c6\u0307 e\u03c6 (1) The above equation yields the relative accelerations as [16] r\u0308 \u2212 r\u03c6\u03072 \u2212 r\u03b8\u03072 cos2 \u03c6 = wr r\u03b8\u0308 cos\u03c6 + 2r\u0307\u03b8\u0307 cos\u03c6 \u2212 2r\u0307\u03c6\u0307\u03b8\u0307 sin\u03c6 = w\u03b8 \u2212 u\u03b8 r\u03c6\u0308 + 2r\u0307\u03c6\u0307 + r\u03b82 cos\u03c6 sin\u03c6 = w\u03c6 \u2212 u\u03c6 (2) where the relative distance r, the pitch line-of-sight angle \u03c6 and the yaw line-of-sight angle \u03b8 , can be measured by the homing sensor: seeker, in terminal phase" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001875_978-4-431-54595-8-Figure6.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001875_978-4-431-54595-8-Figure6.1-1.png", "caption": "Fig. 6.1 The model of attention synchronization", "texts": [ " We believe that a robot should react to a human listener\u2019s behavior and simultaneously generate a situation where the human listener feels comfortable in accepting the robot\u2019s attention-drawing behavior. To build such a robot, we distinguish three processes: attention synchronization, context focus, and believability establishment. \u2022 Attention synchronization This process provides the interacting person with a feeling that the robot is attending to her pointing behavior. The flow of attention synchronization is shown in Fig. 6.1. It consists of three subprocesses: (a) Reaction to the initiation of pointing When the listener notices the initiation of the speaker\u2019s attention-drawing behavior, the listener immediately starts following it with her gaze. (b) Prediction of indicated object Next, the listener estimates the intention of the speaker\u2019s attention-drawing behavior and predicts what object the speaker intends to indicate. The listener often starts looking at the object before the speaker\u2019s pointing motion is finished", " The robot engages in humanlike greetings, reports its \u201cexperience\u201d with products in shops, and tries to establish a good relationship (or rapport) with customers. This is very different from the master-slave-type communication where a robot prompts a user to provide a command. 6.3.2.3.2 Guiding Behavior There are two types of behaviors prepared for guiding: route guidance and recommendation. The former is a behavior in which the robot teaches how to get to a destination with utterances and gestures, as shown in Fig. 6.1. The robot points to the first direction and says \u201cplease go that way\u201d with an appropriate reference term chosen by the attention-drawing model (introduced in the earlier part of this chapter). It goes on to say: \u201cAfter that, you will see the shop on your right.\u201d Since the robot knows all of the mall\u2019s shops and facilities (toilets, exits, parking, etc.), it can teach about 134 destinations. In addition, in situations where a user has not decided where to go, we designed recommendation behaviors: the robot makes suggestions on restaurants and shops" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure5-1.png", "caption": "Fig. 5. External teeth type bearing.", "texts": [], "surrounding_texts": [ "The pitch and yaw bearings used in wind turbines have numerous bolt holes on their inner and outer ring, and they are respectively bolted to the hub and blade and the tower and nacelle during normal operation. To test these conditions, these bearings should be bolted to the corresponding frame. Furthermore, pitch and yaw bearings can be of internal or external teeth type, respectively shown in Figs. 4 and 5, depending on the location of the integrated gear teeth. Pitch bearings are typically of the internal teeth type, whereas yaw bearings can be of both types. Therefore, the test rig should be designed so as to be able to test both types of bearings. For internal teeth type pitch bearings, the outer ring is bolted to the hub and remains stationary, whereas the inner ring is bolted to the blade and rotates according to the wind speed. Therefore, the bearing\u2019s inner and outer rings should be divided into loaded and stationary parts to apply test loads to the loaded part while the stationary part is connected to the fixed supporting structure. The test rig should be able to load and rotate the test bearing simultaneously. It is impossible to perform the two functions simultaneously using only a set of bearings because of the structural configuration of the bearings and the test rig. To solve this problem, by using two sets of the same bearings, two rings with gear teeth attached for both bearings are con- nected to each other using an insert plate and rotated by the driving system whereas two rings without gear teeth attached are connected to the fixed and loaded frame of the test rig, respectively. In this case, between the two bearing sets, the bearing positioned lower is in an environment similar to the operational one. Figs. 6 and 7 show the connection methods for internal and external teeth type bearings, respectively. A bearing driving system comprising a driving motor, reduction gearbox, and pinion connected to the gearbox is used to rotate the test bearing in a manner following the actual environment. The driving systems for pitch and yaw bearings are generally called as pitch and yaw drives, respectively. The rotational speed of the bearing can be controlled by changing the gear ratio of the reduction gearbox and the number of pinion teeth." ] }, { "image_filename": "designv11_22_0001673_icrom.2014.6990767-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001673_icrom.2014.6990767-Figure1-1.png", "caption": "Fig. 1.The schematic of a six-DOF model of REMUS.", "texts": [ " The proposed modified SDRE regulates the depth to desired value without steady-state error while the conventional SDRE fails. The rest of the article is organized as follows. A mathematical model of AUV is expressed in Section II. The structure of the controller is presented in Section III. Implementation of the controller on the model is presented in Section IV and the results of simulation are provided in Section V. Finally, conclusions are expressed in Section VI. II. MATHEMATICAL MODEL OF AUV A schematic view of an AUV with related coordinate systems is presented in Fig. 1 to show a six-DOF AUV model. 978-1-4799-6743-8/14/$31.00 \u00a92014 IEEE 001 The two reference frames are applied to the model: earthfixed frame and body-fixed frame. Descriptions of the parameters are expressed in Table I. The Attitude to present the kinematics of AUV\u2019s model in the global reference frame is based on Euler angles. where 16\u00d7\u211c\u2208\u03b7 is the vector of position and attitude of a vehicle in the inertial frame, 16\u00d7\u211c\u2208\u03bd is the vector of the linear and angular velocities of a vehicle in a fixed-body frame, [ ]Twvu=1\u03bd and [ ]Trqp=2\u03bd " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003261_s10514-020-09959-0-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003261_s10514-020-09959-0-Figure3-1.png", "caption": "Fig. 3 Husky (a) and Hunter (b), the skid-steered wheeled robots used in the experiments", "texts": [ " Then the mean value, denoted by \u03b2\u0303, can be calculated from the individual \u03b2 values using \u03b2\u0303 = 1 N N\u2211 i=1 \u03b2[i] = 1 N N\u2211 i=1 ( uL [i] \u2212 uR[i] w[i] ) (11) where [i] indicates the index of the experiment, and N indicates the number of experiments performed. However, if a calibration sequence is used, the estimated \u03b2 value will be valid for just the corresponding mass distribution. In this work \u03b2 was found using both the mean value, and the proposed approach and the values are compared. In order to verify the kinematicmodel developed in this paper, two skid-steered robotswere used. The firstwas a customized Husky A200 from Clearpath (Fig. 3a), and the second was a robot, nicknamed Hunter, built in our laboratory as part of our continuing research on skid-steered platforms (Fig. 3b). Husky has an almost square base, whose dimensions b and a are 0.55m and 0.51m respectively. The original robot was extendedwith a sensor bridge, and a2degrees of freedomarm increasing its weight from 50kg to approximately 90kg. The robot is equipped with an X-Sens MTi 300 IMU, which was used to measure the rotational speed of the robot. The IMU is a state of the art MEMS IMU with very low bias stability, namely 10\u25e6/h, whichmeans a 10\u25e6 error in the yaw angle if the robot stays stationary for 1 h" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003519_s10846-021-01359-5-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003519_s10846-021-01359-5-Figure9-1.png", "caption": "Fig. 9 An operator remotely maneuvering using two analog joysticks(left for orientation and right for translation control) the leader robot through the obstacles as per visual feedback", "texts": [ " Additionally, flange couplings are custom-designed for the wheels and motors, and machined through lathe operations. The upper platform is a turntable mechanism containing thrust bearings and 3D printed mounts. Furthermore, clamping jaws are provided on the turntable to hold the payload intact, as shown in Fig. 4. Figure 5 shows the control architecture of the leader robot. It consists of a Wi-Fi development kit called as Wemos D1 mini running at 80MHz with ESP8266 chip. The leader robot is remotely controlled from a mobile phone using Blynk App, as seen in Fig. 9. The app contains two analog joysticks which control the translation and rotation of the robot, respectively. Wemos, containing a Wi-Fi module also has a few input-output pins. These are used to signal two channel motor drivers powered by a Li-Po batter of 11.4V. A power bank is used to power the development kit. Therefore, based on the user\u2019s input through the analog sticks, the leader robot moves in the same direction with given velocity using Eq. 2. Four 78 kg-cm rectangular gear box motors are used for this purpose" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003301_j.matpr.2020.12.1110-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003301_j.matpr.2020.12.1110-Figure6-1.png", "caption": "Fig. 6. Stress and strain for magnesium alloy.", "texts": [], "surrounding_texts": [ "The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper." ] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.37-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.37-1.png", "caption": "FIGURE 2.37 Schematic layout of the brush tire model.", "texts": [ " A typical example is when one needs to complete verification analyses using tire data for which only plots of pure slip characteristics are known (e.g., derived from a paper), or if only pure slip tests are available. Such plots or test results can easily be transferred (or well estimated) in Magic Formula parameters for pure slip (see Appendix 7). The expressions (2.73a) and (2.73b) will allow approximate combined slip analysis. We verified this approximation using the same data as used in the previous plot, and for a wheel load of 6000 [N] (see Figure 2.37). This figure corresponds to the polar diagram in Figure 2.30. The behavior for small and large brake slip is quite good. For intermediate brake slip, the brake force is underestimated for the set of Magic Formula data from Appendix 6. Also, observe the symmetry in the approximation based on Eqs. (2.73a) and (2.73b). Changing the sign of the slip angle will only lead to a sign change in Fy without any effect on Fx. The Magic Formula allows a lack of symmetry, as shown in Figure 2.36. In this section, we shall present the theory of steady-state slip with the aid of some simple physical models: \u2022 The brush-type tire model \u2022 The brush-string model In all cases, it is assumed that the properties of the tire can be described by averaging the local behavior over the tire width, which means that the tire is replaced by a disk of zero width" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001046_1.4026442-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001046_1.4026442-Figure6-1.png", "caption": "Fig. 6 Locations from which miniature compression specimens were extracted. (a) Schematic showing the plane definitions and RCF affected regions in a ball, (b) Micrograph of the RCF affected region in meridial section after etching, (c) Micrograph of the RCF affected region in equatorial section after etching, and (d) Miniature compression specimens extracted from the RCF affected zone of meridial section.", "texts": [ " During RCF, only a small material volume undergoes fatigue loading over hundreds of millions of cycles [13,14,16,18,24,25]. The evolution of the constitu- tive response of this region is of critical interest to the bearing materials community with respect to endurance limit and lifing. 3.1 Specimen Extraction. Three 28.6 mm (1.125.) diameter M50 steel balls were received from Pratt & Whitney Inc., East Hartford, CT. These balls had undergone RCF over several hundred million cycles under the conditions shown in Table 1. Only a small region below the contact surface is affected by the RCF; see Fig. 6(a). To locate and reveal the subsurface regions affected by the RCF, these balls were cut in half, perpendicular to the rolling direction, using EDM. This plane, which is perpendicular to the rolling direction, is called the equatorial plane, and the rolling plane is called the meridial plane; see Fig. 6(a). The EDM cut surfaces were polished using the standard metallurgical practices and then a final polishing step was performed using chemicalmechanical polishing (CMP) with a mixture of 0.05 lm colloidal silica and alumina with a pH of 8.5. This procedure slowly Table 1 RCF test conditions and results of compression tests Ball ID Test conditions (Max rHz, temp., and no. of cycles) Specimen ID Yield strength (0.2%) MPa Hardening coefficient, K MPa Hardening exponent, n Ball 2 virgin N/A #1 2620 4850 0", " The RCF affected zone appears as a bright elliptical region on the equatorial surface (see Figs. 6(a) and 6(c)) and as a circumferential ring on the meridial surface (see Figs. 6(a), 6(b), and 6(d)). The size of this region varies depending on the RCF conditions, but the typical size of the major and minor axes of the elliptical region were 3500 lm and 600 lm, respectively. To extract compression test specimens from the small RCF affected zone, a 2 mm thick disk containing the RCF zone was EDM cut along two sections parallel to the meridial plane; see Fig. 6(d). Specimens were then extracted from the RCF affected zone of this disk. In Figs. 6(a) and 6(b), the dotted circles represent the specimens that are to be extracted, while an actual meridial section is shown in Fig. 6(d). Specimens were also extracted from the regions away from the RCF affected zone, which are composed of virgin material. The resulting specimens have EDM damage and also do not meet the previously set requirements for flatness, parallelism, and perpendicularity for a valid compression test. Therefore, these specimens must be further prepared. First, the specimens were manually polished to remove EDM damage from the cylindrical surface. Then, the specimens were finely polished on the loading surfaces to meet the final specifications for flatness and parallelism using a novel approach where gauge blocks and WC tabs were used, as shown in Fig", " Because the global unloading response may be considered elastic [26] unloading after the initial yield provides an accurate measure of elastic modulus. The average of the two unloading steps was taken as the measured modulus. From the force-deflection measurements, the true stress-true strain curve of the specimen is obtained. Typical stress-strain curves are shown in Fig. 11. For clarity, the unloading phase is shown only for one of the curves. For all the specimens, the yield strength was measured at a 0.2% offset. 3.4 Comparison of Virgin and RCF Specimen Response. M50 bearing steel specimens extracted far from the RCF regions (see Fig. 6(a)) were tested to verify the validity of the miniature compression test method described above. From the stress-strain curves provided in Fig. 11, it is noted that the flow curves for the two virgin specimens extracted from two different balls are almost identical. The average measured unloading modulus on each curve was 204.5 GPa, whereas the modulus for M50 reported in literature is 203 GPa [27\u201329]. Similarly, the measured yield strength for virgin M50 was 2610 MPa, whereas the yield strength reported in literature is 2620 MPa [27,29]. Both of these values are within 0.5% of the literature values leading to the conclusion that the procedure followed for specimen extraction and preparation is a valid method for miniature compression testing. After verifying the method by testing the virgin M50 material, specimens extracted from the RCF affected zone (see Fig. 6) were tested. Three balls of diameter 28.6 mm (1.125 in.) diameter, tested under varying RCF conditions by Pratt and Whitney, East Hartford, CT, were used to obtain the constitutive response of the RCF affected material. The test conditions for each of the balls are illustrated in Table 1. Stress-strain curves from one representative specimen from each ball are plotted in Fig. 11. The elastic moduli of all of the specimens plotted are similar, but there is a distinct increase in the yield strength of the RCF affected regions depending on the test conditions of the ball" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003601_tmag.2021.3081799-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003601_tmag.2021.3081799-Figure3-1.png", "caption": "Fig. 3. Basic dimensions of the proposed FM-TFLG. (a) Sectional view. (b) Side view.", "texts": [ " The Fourier series expansion of armature reaction MMF can be expressed as, 1 1 3 , sin 2 2 2 1 sin 4 q q rms q S F t M NI S M q q (5) where S is a constant in armature reaction MMF, qM is a coefficient of armature reaction MMF determined by q , N is the number of alternating current coil turns, 1 is the half of arc between PMs, q is number of slots per phase per pole. 4 , 3 1 = 4 , 3 1 0, 3 t t t t q N t q r q N t q r q r (6) where is a constant and r is a positive integer. III. OPTIMIZATION OF DIMENSION PARAMETERS The dimension parameters of the proposed FM-TFLG are given in Fig. 3. The definitions of these parameters are listed in Table I. Then, in the next section, the dimension parameters will be optimized to achieve higher thrust / power density and lower total harmonic distortion (THD) of back-EMF. TABLE I SOME DIMENSION PARAMETERS OF THE GENERATOR Symbol Parameter Symbol Parameter isir Inner stator inner radius osor Outer stator outer radius isyr Inner stator yoke radius pm PM arc issr Inner stator shoe radius in Translator inner arc isor Inner stator outer radius out Translator outer arc tir Translator inner radius sh Outer stator shoe arc tor Translator outer radius tw Outer stator tooth width osir Outer stator inner radius yh Outer stator yoke height ossr Outer stator shoe radius tl Translator length osyr Outer stator yoke radius twist Translator twist arc Authorized licensed use limited to: Carleton University" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001646_j.procir.2016.01.078-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001646_j.procir.2016.01.078-Figure3-1.png", "caption": "Fig. 3. (a) Isometric view of tracheal implant model; (b) CT images of the porcine trachea used to determine the elliptic dimensions of the (c) implant model.", "texts": [ " Optimized results from GA. Parameter Value wE1 =1 wE2 =1 w =1 Vf of CNT 9\u00d710-6 Vf of Ag Particles 1.6\u00d710-5 Vf of PDMS 0.999 E1 9.8179MPa E2 14.99MPa 965kg/m3 wE1 =12 wE2 =5 w =1 Vf of CNT 2.01\u00d710-5 Vf of Ag Particles 3.3\u00d710-5 Vf of PDMS 0.999 E1 17.998MPa E2 29.72MPa 965kg/m3 The optimized material composition is then input into FEM software for evaluation. Briefly, CT images of the subject\u2019s trachea were first acquired and the geometrical design of the implant model based on the medical images (see Fig. 3). 3D reconstruction of the patient\u2019s tracheal was also performed and both tracheal implant model and tracheal model were integrated in the FEM software for simulation (see Fig. 4). Through this analysis, we are able to compare the stress concentration within the implant and surrounding tissues and improve the design parameters. 4. Discussion The proposed method of rule-based intelligent design has be evaluated and demonstrated in the design and optimization of a patient specific tracheal implant model" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002379_0954406220917424-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002379_0954406220917424-Figure4-1.png", "caption": "Figure 4. Relative displacement between the members of the single stage.", "texts": [ " In Figure 3, xi, yi, zi and uix, uiy, uiz (i\u00bc s, c, r, n) are the elastic deformations and torsional deformations around the axes of each component, respectively. kix, kiy and cix, ciy are the bearing stiffness and damping of the corresponding component in the x, y directions, respectively. kju and cju (j\u00bc s, c, r) represent the torsional stiffness and damping of s, c, and r, respectively. krn, ksn and crn, csn denote the mesh stiffness and damping of the nth sun\u2013planet (s\u2013p) and planet\u2013ring (p\u2013r) gear pairs, respectively. The relative displacement between members of stage n is shown in Figure 4. The following equations can be derived from Figure 4. 1. Relative displacement along the LOA for the sun\u2013 planet mesh sn \u00bc xs sin sn cos ys cos sn cos zs sin usz cos xn sin 0 ts cos \u00fe yn cos 0 ts cos \u00fe zn sin unz cos usx sin n sin \u00fe usy cos n sin \u00fe uny sin = cos 0 ts \u00fe esn \u00f09\u00de 2. Relative displacement along the LOA for the ring\u2013 planet mesh rn \u00bc xr sin rn cos yr cos rn cos \u00fe zr sin urz cos \u00fe xn sin 0 tr cos \u00fe yn cos 0 tr cos zn sin \u00fe unz cos \u00fe urx sin n sin ury cos n sin \u00fe uny sin = cos 0 tr \u00fe ern \u00f010\u00de where 0ts and 0 tr represent the transverse working pressure angle of the s\u2013p and p\u2013r gear pairs, respectively, sn \u00bc 0 ts \u00fe n and rn \u00bc 0 tr n denote the working pressure angle of the external and internal meshing pairs, respectively, n \u00bc 2 \u00f0n 1\u00de=N is the position angle of the nth planet gear, and esn and ern are the static transmission error of the s\u2013p and p\u2013r pairs, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002058_1.4035203-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002058_1.4035203-Figure3-1.png", "caption": "Fig. 3 Schematic of the stern support", "texts": [ "url=/data/journals/jvacek/935929/ on 02/09/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use and the slim shaft was supported by three supports. In the stern support, as shown in Fig. 2, a water-lubricated rubber bearing was particularly used to investigate vibration associated with the bearing friction, and four beams were connected to the bearing housing and pull the bearing radially to make it located firmly at the center. The connection between the shaft and the support is illustrated in Fig. 3. As can be seen, one end of the beam was hinged with the bearing housing and the other end, pretightened longitudinally by the force FP, was fixed onto the foundation. A piezoelectric force sensor (HD-YB-311), as shown in Figs. 1 and 3, was inserted in series between the outside-end of each horizontal beam of the stern support and the foundation. 2.2 Experimental Results. Experiments on friction-induced vibrations in the stern support were conducted, and the dynamic reaction forces transmitted from the horizontal beams in the stern support to the foundation were measured at different shaft speeds", " To find out the associated unstable modes and derive modal parameters for further analysis, a finite element model of the flexibly supported shafting system submerged in water, as shown in Fig. 9, was built with the software ABAQUS. Parameters of the model are listed in Table 1. The water and the shaft are modeled with 506,496 solid elements, and the supports are modeled with 9472 shell elements. The supports and the shaft are connected by springs. To consider the influence of water, the nodes at the surfaces between the shaft and supports are coupled with the associated nodes of water, which have a nonreflecting boundary. The outside-ends of support beams as shown in Fig. 3 are fixed, and the pretightened force Fp is 1700 N. The first six modes of the stern support are illustrated in Fig. 10, and the plotted shapes include undeformed and deformed ones in order to show the changes between them. It can be seen evidently that the first, third, fifth, and sixth modes are torsional modes because the two circles associated, respectively, with the undeformed and deformed shapes are apart, and especially, the fifth mode of the stern support is associated with instability. In Secs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003727_01423312211021294-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003727_01423312211021294-Figure1-1.png", "caption": "Figure 1. Diagram remote control of unmanned helicopter.", "texts": [ " (3) With 3-DOF helicopter as the test bench, comparative experiments are conducted, in which both wind gust and communication delay tests are carried out. Also, the controller performances under the different wind velocities and random time varying delays are demonstrated separately. The following parts of this paper are organized as follows. The 3-DOF helicopter model and problem description are introduced in Section 2. In Section 3, design of the robust attitude controller method is presented. In Section 4, the experimental results are displayed. Sections 5 states the conclusions. Figure 1 shows the diagram of remote control for 3-DOF unmanned helicopter that has been widely utilized because it possesses numerous crucial practical features of unmanned helicopter, including strong coupling, parameter uncertainties and nonlinearity. It is convenient to carry out practical experiments to validate the control performance and effectiveness of predesigned flight control based on the platform. Due to the use of wireless network in remote control of unmanned helicopter, network-induced time-varying delays would be appeared in control system (Liu et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001638_etep.2165-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001638_etep.2165-Figure1-1.png", "caption": "Figure 1. Schematics of two samples of three-phase SynRMs with (a) four and (b) two poles (c) supplying and connection diagram.", "texts": [ " Noticing that the SynRM has been less investigated for studying eccentricity fault diagnosis, this research is focused on studying of SynRM\u2019s behavior in healthy and faulty condition using a mathematical modeling. Neglecting core permeability, performance analysis for torque\u2019s ripple computation of four samples of three-phase SynRMs has been considered in this research as follows: \u2022 Case 1 4-pole-60-slot SynRM \u2022 Case 2: 2-pole-30-slot SynRM \u2022 Case 3: 2-pole-24-slot SynRM \u2022 Case 4: 4-pole-36-slot SynRM Figure 1 shows the structure of cases 1 and 2. As shown in Figure 1, reference of stator circumambient (\u03b8s=0) corresponds with the center of winding \u201cA\u201d [27] and the reference of rotor circumambient Copyright \u00a9 2015 John Wiley & Sons, Ltd. Int. Trans. Electr. Energ. Syst. (2015) DOI: 10.1002/etep (\u03b8r=0) is according to maximum air gap length [26,28]. Considering Ns turn for each pole per phase, Ns/5 turns will be achieved for each slot. Moreover, as shown in Figure 1, an inductance has been considered for star connection and Ln\u2192\u221e can model a float star connection. Machine\u2019s parameters have been shown in Figure 1. Now, there are some parts that should be modeled for dynamic modeling. (A) Distributed winding Sinusoidal winding function is an ideal property of machine that cannot be assumed in an accurate model. In fact, limited number of slots is one of the real factors that lead to non-sinusoidal winding, turn, and magnetomotive force functions. Considering ns slots and Ns turns per pole in a three-phase P-pole machine denotes ns/3P slots per pole and 3P.Ns/ns turns per slot. Now, details of the turn function for the phase A and stator\u2019s turn functions for a sample machine with five slots per pole can be considered as Figure 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002620_s13198-020-01019-1-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002620_s13198-020-01019-1-Figure5-1.png", "caption": "Fig. 5 PMSM four pairs of poles Fig. 6 Three-phase winding with faulty internal winding in phase a", "texts": [ " Also, permanent magnet synchronous motors have more efficiency and power than other motors. However, their application with high power is not wide, and this relates to the ability to use permanent magnetic materials, the cost, and the ability to manufacture these materials. Due to the design of the rotor, the PMSM poles can be designed in a round or salient pole. Round pole motors have a cylindrical rotor and thus have even air spacing. Salient pole motors have variable air spacing. A model of a permanent magnet synchronous motor with four poles is depicted in Fig. 5. Avoiding faults in engineering systems is not easy. This is while faults can have dangerous consequences. Troubleshooting and effective diagnostics can improve the reliability of the system and prevent costly maintenance and repairs. The identification of faults in permanent magnet synchronous motors is of great importance because of the critical applications in the battleships and aerospace. In the powerful design of a fault detection system, the first step is to extract the appropriate knowledge through the provision of a consistent model that can simulate the system\u2019s entire behavior" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001045_cdc.2014.7039508-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001045_cdc.2014.7039508-Figure1-1.png", "caption": "Fig. 1. Schematic representation of the vehicles", "texts": [ " The navigation law to drive back the UAV to the launch base is then considered in Section IV. Simulation and experimental results are finally given in Section V before terminating with concluding remarks. In the subsequent sections, two ground units are considered and are composed of two vehicles V1 and V2 each equipped with an antenna. The drone, considered as a quadrotor helicopter UAV, is assumed to have a constant altitude h during the mission. The UAV is also equipped with one antenna. A schematic representation of the vehicles is depicted in Figure 1 where V1 and V2 denote the ground vehicles and P1 and P2 are the signals strength as measured by the UAV\u2019s on-board antenna. Seen from above, xy represents the body frame of the UAV. \u03b8i represents the angle of arrival of the signal from the ith vehicle in the xy plane. Several methods may be employed to determine the angle of arrival of signals. In [13], a proof-of-concept device and a method are described to estimate the direction-of-arrival (DoA) of a radio signal by a receiver. The device estimates the DoA by identifying the peak of received signal strength indicator measurements using an actuated parabolic reflector", " (x(t), y(t)) are the coordinates of the drone and dgi (t) is the ground distance between the drone and the ith vehicle, i.e. dgi (t) = \u221a (xi \u2212 x(t))2 + (yi \u2212 y(t))2. A side view of the vehicles is shown in Figure 3 where the cross-section plane is the vertical plane containing the UAV and the ith vehicle. This figure shows the height h of the UAV with respect to the ground vehicles as well as the distance di from the UAV to the ith vehicle. The distance di is given as function of the ground distance dgi and height h as d2 i = (dgi )2 + h2. The explicit model for the signals strength Pi (i = 1, 2) shown in Figure 1 is presented in the subsequent section. The propagation model for the signals strength Pi as measured by the UAV\u2019s on-board antenna is given as: Pi = Ti +GTxi +GRx\u2212 PLi (1) where Ti is the transmission power, GTxi and GRx are the gains of the transmitting and receiving antennas respectively, and PLi is the path-loss. Various path-loss models exist in the literature and they typically have the following form [16]: PL = A log10(d) +B + C log10 ( fc 5 ) +X (2) where d is the distance between the transmitter and the receiver" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003610_s11665-021-05762-9-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003610_s11665-021-05762-9-Figure5-1.png", "caption": "Fig. 5 Computational domain", "texts": [ " Dilution definitions can determine a proper layer in terms of geometrical aspects, as shown in Fig. 4, through optical microscopy. Dilution (%) can be calculated using the following Eq 2 (Ref 11), where, d: height of the mixing area (depth) and h: height of the layer area. Dilution \u00bc d d \u00fe h 100% \u00f0Eq 2\u00de The gas flow during the deposition was analyzed through a CFD model simulation. The computational domain was created using only half of the nozzle\u2019s annular geometry taking advantage of its symmetry in the analysis. This decision significantly reduces the computational cost. Figure 5 shown the employed computational domain where inner green presents gas nozzle inlet; middle green the gas shield inlet; outer green the gas carrier inlet; gray the nozzle walls; red the outlet pressure; and finally, the substrate represented by the bottom. The numerical analysis was focused on the flow behavior of argon between the volumetric flow inputs nozzle and the substrate in order to analyze fluid behavior with different input values The simplifying assumption is the following: \u2022 The gas is treated as steady-state turbulent flow \u2022 Flow with constant velocity distribution in the inlet boundaries \u2022 Boundaries inlet flow have the same gas source \u2022 Constant density \u2022 Pressure gage equal zero \u2022 Slip condition at the walls The whole complexity of the nozzle geometry and size of the computational domain was maintained according to the original to get more accurate results" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002709_fuzz48607.2020.9177761-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002709_fuzz48607.2020.9177761-Figure5-1.png", "caption": "Fig. 5. A three-link robot", "texts": [], "surrounding_texts": [ "4) Weight memory: The weight is define as:\n1 k\nT\nj j kj n jw w w \u22ef \u22efw (11)\nwhere kjw is the output weight.\n1 j mw w w \u22ef \u22efW (12)\n5) Output: The j-th output of MGMFFC is calculated as\n1 1 1\n, k kn n n synthesis previous current\nj kj k kj k k k ik\nk k i u w b w b \n \n (13)\nThe MGMFFC output is rewritten via a vector form:\n11 12 1 1\n21 22 2 2\n1 2\nk\nk\nkk\nsynthesis n\nsynthesis n\nsynthesis\nnm m mn\nw w w b\nw w w b\nbw w w \u22ef \u22ef \u22ee\u22ee \u22ee \u22f1 \u22ee \u22ef u Wb (14)\nB. Adaptive Learning Laws for MGMFFC\nTaking the derivative of ( ( ), )S t ts e and using (1), yields\n( )( ( ), ) ( ) ( )n T S St t t t \u027as e e e\n( )( ( )) ( ( )) ( ) ( ) ( ( )) ( )n T ref S t t t t t t G H u y l e (15) Insert (7) into (15) and multiply both sides by ( ( ), )T\nS t ts e ,\ngives:\n ( ( ), ) ( ( ), ) ( ( ), ) ( ( )) ( ( ), ) ( ( )) ( ) ( )\nT T\nS S S\nT\nS RBMGMFFC\nt t t t t t t\nt t t t t\n \n \n\u027as e s e s e G\ns e H u u\n\n ( )( ( ), ) ( ) ( ( )) ( )T n T\nS ref St t t t t s e y l e (16)\nwhere 1 ( )( ) ( ( )) [ ( ) ( ( )) ( ( )) ( )]n T\nreMGMFF fC S t t t t t t u H y G l e (17) Define a Lyapunov cost function as:\n 1 ( ), ( ), ( ),\n2\nT\nS S SV t t t t t ts e s e s e (18)\n ( ), ( ), ( ),T S S S V t t t t t t \u027a \u027as e s e s e . The aim is to minimize ( ( ), ) ( ( ), )T\nS St t t t\u027as e s e for reaching fast convergence\nof s . The gradient descent technique is then used so the\nparameters are updated through the online adaptive laws as follows:\n ( ), ( ), ( ), ( ), \u02c6 T T jS S S S\nkj w w\nkj j kj\nut t t t t t t t w\nw u w \n \n \n\u027a \u027a \u027a s e s e s e s e\n0( ) synthesis w j j ks t g b (19)\n ( ), ( ), ( ), ( ), \u02c6 T T current jS S S S ik\nik m m current\nik j ik ik\nut t t t t t t t m\nm u m\n \n\n \n \n\u027a \u027a \u027a s e s e s e s e\n0 2 1\n2( ) ( )\nn current i ik\nm j j kj k ik\ni ik\nI m s t g w\nv \n\n (20)\n ( ), ( ), ( ), ( ), \u02c6 T T current jS S S S ik\nik v v current\nik j ik ik\nut t t t t t t t v\nv u v\n \n\n \n \n\u027a \u027a \u027a s e s e s e s e\n2\n0 3 1\n2( ) ( )\nn current i ik\nv j j kj k ik\ni ik\nI m s t g w\nv \n\n \n (21)\n ( ), ( ), ( ), ( ), \u02c6 \u02c6 T T previous jS S S S k k previous\nj k kk\nut t t t t t t t b\nu b \n \n \n\u027a \u027a\u027a s e s e s e s e\n2 0( ) previous j j kj k ks t g w b (22)\n ( ), ( ), ( ), ( ), \u02c6 \u02c6 T T current jS S S S ik k current\nj ik kk\nut t t t t t t t\nu \n \n \n \n \n\u027a \u027a\u027a s e s e s e s e\n2\n0\n1\n( ) n current\nj j kj k ik\ni s t g w (23) where ju is the jth element of ,u and m , , w , and\n are respectively learning rates for means, variances, weights, and adaptive gains for the current and previous states.\nC. Robust controller design\nThis study uses a sgn(.) function robust compensation\ncontroller to cover the approximation error:\n 1 \u02c6 ( ( )) sgn( ( ), )RB St t t u H s e (24)\nThe error bound is updated online by:\n\u02c6 \u027a s (25)\nwhere are positive learning rates.\nProof: Define a Lyapunov function as\n 2\n( ( ), , ) 2 2\nT\nSV t t \n \n \n\u0276 \u0276 s s s e , (26)\nDifferentiating (26) with respect to time then using (15) and\n(24), gives:\n ( ( ), , ) T\nSV t t \n \n \n\u027a\u0276 \u0276\u027a \u0276 \u027as e s s\n \u02c6 sgn( )T\n\n \n \n\u027a\u0276 \u0276 s \u03b5 s\n \u02c6T\n\n \n \n\u027a\u0276 \u0276 s \u03b5 s . (27)\nAuthorized licensed use limited to: SUNY AT STONY BROOK. Downloaded on October 05,2020 at 04:28:06 UTC from IEEE Xplore. Restrictions apply.", "The approximation error \u03b5 is the minimum reconstructed error\nbetween *\nID ( )tu and ( )MGMFFC tu , and it is assumed to be\nbounded by 0 \u03b5 .\nIf the adaptive law of the error bound is selected as\n\u02c6 \u027a\u027a\u0276 s , (28)\nthen (27) becomes:\n \u02c6 \u02c6( ( ), , ) T SV t t \u027a \u0276s e s \u03b5 s s\n T s \u03b5 s \u03b5 s s\n 0 \u03b5 s . (29)\nSince ( ( ), , )SV t t \u027a \u0276s e is negative semi-definite that is\n ( ( ), , )SV t t \u027a \u0276s e ( (0), (0))V \u027a \u0276s , it implies that s and\n\u0276 are bounded. Let function \u03b5 s\n ( ( ), , )SV t t \u027a \u0276\u03b5 s s e , and integrate with\nrespect to time, yields\n0\n( ) ( (0), (0)) ( , ) t\nd V V s s\u027a \u027a\u0276 \u0276 . (30)\nBecause ( (0), (0))V s\u027a \u0276 is bounded, and ( , )V s\u027a \u0276 is non-\nincreasing and bounded, the following result is obtained:\n0\nlim ( ) t\nt d . (31)\n\u027a is bounded, so lim ( ) 0t t . Then, 0s when t . Finally, the MGMFFC is asymptotically stable.\nTherefore, the proof is complete.\nThe dynamic equation is taken in [14] as:\n( ) ( , ) ( ) d\n \u027a\u027a \u027a \u027aq q q q q g q \u03c4 \u03c4 , (32)\nwhere\n1 4 2 5 23 2 4 2 6 3 3 5 23 6 3\n4 2 5 23 2 6 3 3 6 3\n5 23 6 3 3\n2 2 2 1 0 0\n( ) 2 2 1 1 0\n2 1 1 1\nd d c d c d d c d c d d c d c\nd c d c d d c d d c\nd c d c d q ,\n5 23 5 234 2 5 23 2 32 2\n4 2 5 232 2\n6 3 5 23 6 3 5 233 3 3 1\n5 23 6 33 3\n4 2 5 23 6 3 6 31 1 1 2\n6 3 4 2 5 233 1 1 3( , )\nq d s q d sq d s q d s q d s q d s q d s q d s q d s q d s q d s q d s\nq d s q d s q d s q d s\nq d s q d s q d s q d\n \n \n\u027a \u027a\u027a \u027a \u027a \u027a \u027a \u027a \u027a \u027a \u027a \u027a\n\u027a \u027a \u027a \u027a\n\u027a \u027a \u027a \u027a \u027aq q 6 3 6 3 1 2 3\n5 23 6 3 6 3 6 31 1 2 1 2\n( )\n( ) 0\ns d s q q q\nq d s q d s q d s d s q q\n \u027a \u027a \u027a \u027a \u027a \u027a \u027a \u027a ,\n1 1 1 1 2 12 1 1 2 12 3 123\n1\n2 12 2 12 3 123 2\n3\n3 123\n1 1 1\n2 2 2\n1 1 ( ) 0 2\n2 2\n1 0 0\n2\na c a c a c a c a c a c m g\na c a c a c m g\nm g a c g q , and\n0.2 (2 )\n0.1cos(2 )\n0.1 ( )\nd\nsin t\nt\nsin t \u03c4 .\nAll system parameters for the three-link robot are listed in Table 1. Definitions of variables in (32) are described in Table 2 as\nAuthorized licensed use limited to: SUNY AT STONY BROOK. Downloaded on October 05,2020 at 04:28:06 UTC from IEEE Xplore. Restrictions apply.", "follows:\nvariances are 1.1ikv , for 1, 2, , 10k \u22ef and 1, 2 and 3i . The learning rates for MGMFFC are 1, 1,I P 0.05,w m 0.01, eI m D 0.1, 0.1, eD 1, and 0.1 . The initial value of\n0 0and are set randomly between -1 and 1. In order to show\nthe effectiveness of the MGMFFC, the TFLFCM (TOPSIS Function-link CMAC) [15] is also used for the three-link robot to compare their performance. The angular trajectories and the joint velocities are respectively plotted in Figs. 6 and 8 (a)-(c). Enlarge of angular trajectories and joint velocities are shown in Figs. 7 (a)-(c) and Figs. 9 (a)-(c), respectively. The control efforts and their enlargements are displayed in Figs. 10 and 11 (a)-(c). The tracking errors and the enlargement are plotted in Figs. 12 and 13 (a)-(c). Finally, two adaptive gains and of\nthe MGMFFC are displayed in Fig. 14. The simulation results indicate that the proposed MGMFFC achieves excellent control performance under external disturbance. The tracking errors for three links converge quickly to zero (see Figs. 12 and 13 (a)(c)). The tracking angular positions of three links follow the reference angular positions well (see Figs. 6 and 7 (a)-(c)). The response of the proposed MGMFFC is fast (see Fig. 10 and 11 (a)-(c)). The proposed MGMFFC achieves a favorable tracking response when two adaptive prediction gains and are\nadjusted online as shown in Fig. 14. The total RMSE (root mean square error) for the TFLFCM and the proposed MGMFFC are measured in Table 3. In summary, Table 3 and the simulation results confirm that the proposed MGMFFC achieves better tracking performance with quicker convergence and smaller tracking error than for TFLFCM.\nAuthorized licensed use limited to: SUNY AT STONY BROOK. Downloaded on October 05,2020 at 04:28:06 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_22_0003009_j.mechmachtheory.2020.104223-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003009_j.mechmachtheory.2020.104223-Figure3-1.png", "caption": "Fig. 3. Free-body diagram of the moving platform.", "texts": [ " According to the vector closure principle, the inverse kinematic equations can be written as \u03c1 j = a j + Qb j + h , (1) where \u03c1 j is the vector from point A j to point B j along the jth cable, a j is the vector from point A j to origin O, b j represents the vector oB j in the reference F 1 . h is the vector of the rigid link, namely the vector from O to o, as shown in Fig. 2 . Q is the rotation matrix from the fixed frame to the moving frame. The effective cable length \u03c1 j is given by \u03c1 j = \u221a (a j + Qb j + h ) T (a j + Qb j + h ) . (2) e j = \u03c1 j /\u03c1 j is defined as the unit vector in the direction of the jth cable, oriented from the static platform to the moving platform. Free-body diagram of the moving platform is shown in Fig. 3 . Then the dynamic model is built using the NewtonEuler approach (all frictions are ignored), which yields 4 \u2211 j=1 f j + \u03ba + m g = 0 , (3) 4 \u2211 j=1 (Qb j ) \u00d7 f j + Qp \u00d7 (m g ) = Q (I \u02d9 \u03c9 + \u03c9 \u00d7 I \u03c9 ) , (4) where m is the mass of the moving platform. I is the inertia tensor per unit mass of the platform relative to the moving reference frame. \u03c9 and \u02d9 \u03c9 are, respectively, the angular velocity and acceleration of the platform with respecting to the moving frame F 1 . \u03ba is the force applied by the rigid link" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001513_j.proeng.2015.12.642-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001513_j.proeng.2015.12.642-Figure1-1.png", "caption": "Fig. 1 Specimen geometries: ruler a), mini-ruler b), tubes c) [2]", "texts": [ " Nomenclature a fiber orientation tensor E Young\u2019s Modulus FO nominal fiber orientation fs fiber share value t thickness Kt stress concentration factor \u03bb biaxiality ratio N, Nexp experimental cycles to failure \u03bd Poisson\u2019s ratio Ncalc estimated cycles to failure R load ratio \u03c1 density S nominal stress Sa nominal axial stress amplitude \u03c3 stress tensor \u03c31,2,3 principle stresses \u03c3a local stress amplitude \u03c3W local alternate fatigue strength \u03c3\u03b3\u03c6 normal stress in the critical plane t, t1,2 specimen thickness Ta nominal shear stress amplitude T scatter index based on the stress amplitude local shear stress a local shear stress amplitude W local alternate shear fatigue strength \u03b3\u03c6 shear stresses in the critical plane 2. Specimens and Experimental Investigation Three different kinds of specimens were tested under fatigue loading with constant amplitude. The main differences between those specimens were the size, the shape, and the manufacturing process resulting in different geometry, surface treatment and the fiber orientation distribution. In Fig. 1 a-c the different specimens are shown: Plain injection molded specimens with a rectangular cross section (\u201crulers\u201d), plain specimens with smaller dimensions milled out of 80x80xt plates (\u201cmini-rulers\u201d) and tubular injection-molded specimens (\u201ctubes\u201d). The relative humidity of the material PA66-GF35 could be described as dry as molded at the beginning of the fatigue experiment. The testing facility did not control the moisture or the temperature. The load ratio for all experiments was R=-1 and all tests were conducted at room temperature. To avoid buckling an anti-buckling device was used for the ruler-type specimens. The tests were carried out by De Monte et. al. [2,3,4] at testing facilities at Bosch and at Fraunhofer LBF in Darmstadt. The rulers have larger dimensions compared to the mini-rulers. Their shape is shown in Fig. 1 a). Moreover the longitudinally injected specimens showed a higher degree of anisotropy caused by a longer melt flow path. For the transversally injected specimens the melt flow path was shorter. This type of specimen was milled out of injection-molded plates (80mm x 80mm x t) in three different angles, see Fig. 1 b). An Angle of the nominal fiber orientation FO=0\u00b0 is equal to a longitudinal fiber alignment and FO=90\u00b0 to a perpendicular fiber orientation. Two different plate thicknesses were considered (t1=1mm, t2=3mm). Injection molded planar specimen exhibit different fiber orientation in the vicinity of the surface compared to the core area. This so-called skin-core effect depends on the melt flow conditions in the cross-section area of the specimens and, as such, is also influenced by the thickness. The skin layer shows a high degree of fiber alignment in flow direction and in the core layer the fibers are predominantly aligned in perpendicular direction [5]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002046_1.4035091-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002046_1.4035091-Figure1-1.png", "caption": "Fig. 1 Configuration of missile system", "texts": [ " They are represented as _a \u00bc KQ Cx sin a\u00fe Cz cos a\u00f0 \u00de tan b p cos a\u00fe r sin a\u00f0 \u00de \u00fe q _b \u00bc KQ Cy cos b Cx cos a cos b Cz sin a sin b r cos a p sin a _p \u00bc IzzIxz \u00fe Ixz Ixx Iyy\u00f0 \u00de IxxIzz I2 xz pq\u00fe Izz IxxIzz I2 xz QSdCl \u00fe Ixz IxxIzz I2 xz QSdCn _q \u00bc 1=Iyy Ixz r2 p2 \u00fe Izz Ixx\u00f0 \u00depr \u00fe QSdCm _r \u00bc I2 xz \u00fe Ixx Ixx Iyy\u00f0 \u00de IxxIzz I2 xz pq\u00fe Ixz IxxIzz I2 xz QSdCl \u00fe Ixx IxxIzz I2 xz QSdCn _/ \u00bc p\u00fe tan h sin /q\u00fe cos /r\u00f0 \u00de _h \u00bc cos /q sin /r _w \u00bc sin /q\u00fe cos /r cos h (1) where a is the angle-of-attack; b is the sideslip angle; p, q, and r are the roll, pitch, and yaw rate about the body axes; and /, h, and w are the roll, pitch, and yaw angles. Figure 1 illustrates the missile system configuration and the definition of these variables. Ixx, Iyy, and Izz are the roll, pitch, and yaw moments of inertia, 041002-2 / Vol. 139, APRIL 2017 Transactions of the ASME Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/journals/jdsmaa/936049/ on 02/15/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use respectively, and are approximated as constants. Q is the dynamic pressure. S and d are the reference area and diameter", "75 and an altitude of 40,000 ft, and thus, the total translational velocity and weight (W\u00bc 227 lb) are assumed to be constant. KQ is a constant and represented as KQ \u00bc \u00f0QSg=WV\u00de. g is the gravity acceleration. The aerodynamic force and moment coefficients (Cx;Cy;Cz; Cm;Cl;Cn) in the preceding equations are complicated nonlinear functions of dynamic pressure, angle-of-attack, sideslip angle, control surface deflection (dp, dq, dr), etc. It is noted that the missile system has four tail fins as shown in Fig. 1. The aileron (dp), elevator (dq), and rudder (dr) deflection angles are effectively generated through intermixing the four fins [13]. Since each fin actuator has a finite bandwidth, the control surface deflection can be modeled as a firstorder system with surface position and rate saturation _dp;q;r\u00f0s\u00de \u00bc \u00bdxa=\u00f0s\u00fe xa\u00de dc\u00f0s\u00de (2) where xa is the actuator bandwidth, and dc is the commanded control surface deflection. A 55 deg maximum deflection magnitude and a 300 deg/s maximum deflection rate are imposed on the actuators of this BTT missile" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001083_1.4027926-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001083_1.4027926-Figure3-1.png", "caption": "Fig. 3 Impulse experiment with compressor rotor experimental rig", "texts": [], "surrounding_texts": [ "4.1 Faults Produced on Ball Bearings. To study the casing responses caused by ball bearing faults, in this study, faults are produced on ball bearings using wire-electrode cutting technology. These faults are produced on the inner ring, outer ring, and ball of the compressor experimental rig\u2019s 6214 ball bearing and the aero-engine experimental rig\u2019s 6205 ball bearing, which are shown in Figs.7 and 8. The ball bearing dimensions are listed in Table 1. The ball bearing faults\u2019 characteristics frequencies can be computed as follows: (1) Outer race foc \u00bc Z 2 1 d D cos a fR (1) (2) Inner race fic \u00bc Z 2 1\u00fe d D cos a fR (2) (3) Rolling ball fbc \u00bc Z 2 D d 1 d D cos a 2 \" # fR (3) (4) Cage fc \u00bc 1 2 1 d D cos a fR (4) 4.2 Faults Experiments of Ball Bearings. The bearing faults experiments are carried out. The test-site photo of the compression rotor rig is shown in Fig. 9, and the measurement points\u2019 explanation is listed in Table 2; the test-site photo of the aeroengine rotor rig is shown in Fig. 10, and the measurement points\u2019 explanation is listed in Table 3. Vibration signals are collected by means of the USB9234 data acquisition card of the NI Company, the 4805 type ICP acceleration sensors of B&K Company are used to pick up the acceleration signals, and the eddy current sensors are used to measure the rotating speeds. The sampling frequency is 10.24 kHz. 4.3 Wavelet Envelope Analysis for Ball Bearing Fault 4.3.1 Basic Principles of Wavelet Envelope Analysis. In this study, a signal analysis to determine ball bearing faults is carried out by means of a wavelet envelope spectrum analysis. Chen [2] provides a reference for the detailed process of this algorithm. The essence of ball bearing fault diagnosis based on the wavelet packet is to take advantage of its bandpass filter characteristics, and to decompose the signals using appropriate wavelet functions so as to obtain an appropriate resonance frequency band. Then, by means of envelope demodulation, low frequency envelope signals that only contain the fault characteristic information are obtained. Its spectrum is the wavelet envelope spectrum, in which the fault characteristic frequencies of the ball bearings can be found out. 4.3.2 Wavelet Envelope Spectrum Analysis of Ball Bearing Fault Signals 4.3.2.1 Experiment Analysis Based on Compressor Rotor Experimental Rig (1) Feature extraction for inner ring faults Figures 11\u201314, respectively denote the time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope spectra of the bearing house\u2019s response and the casing\u2019s response to inner ring faults in the 6124 ball bearing. The experimental rotating speed is 1793 rpm (29.88 Hz). The number of balls is 10, the other ball bearing\u2019s parameters are listed in Table 1, and the inner ring characteristic frequency can be calculated by formula (2) as fic\u00bc 5.8974 29.88\u00bc 176.24 Hz. From Fig. 11, the vibration acceleration response value of the casing is larger than that of the bearing housing, which shows that the impulse vibration of the bearing fault is effectively transmitted to the casing. This result is in full agreement with the experimental result for the impulse response. From Fig. 12, it can be found out that there are many resonance peaks in the frequency spectrum, and the signals are very weak in the low frequency segments, the signal\u2019s frequency spectrum from 0 Hz to 500 Hz are shown in Fig. 13, from which, the characteristic frequency of the inner ring fault and its modulation frequency fr can be Journal of Dynamic Systems, Measurement, and Control NOVEMBER 2014, Vol. 136 / 061009-3 Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use basically seen. The wavelet envelope spectrum is shown in Fig. 14, by comparison with Fig. 11, the characteristic frequency of the inner ring fault and the modulation frequency fr are more distinctly shown in Fig. 14. Obviously, the inner ring fault features can be all extracted from the acceleration signals of the bearing house and the casing by means of the wavelet envelope spectrum analysis and frequency spectrum. (2) Feature extraction of outer ring faults The time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope spectrum of the bearing house responses and casing responses of the 6124 ball bearing outer ring faults are respectively shown in Figs. 15\u201318. The experimental rotating speed is 1826 rpm (30.43 Hz). The number of balls is 11, the other ball bearing\u2019s parameters are listed in Table 1, and the outer ring characteristic frequency can be calculated by formula (1) as foc\u00bc 4.5879 30.43\u00bc 140 Hz. From Fig. 15, the vibration acceleration response value of the casing is larger than that of the bearing housing, which shows that the impulse vibration of the bearing fault is effectively transmitted to the casing. This result is in full agreement with the experimental result for the impulse response. From Fig. 16, it can be found out that there are many resonance peaks in the frequency spectrum, and the Table 1 Ball bearing dimensions (in mm) Type Diameter of inner ring Diameter of outer ring Thick Diameter of ball Pitch diameter" ] }, { "image_filename": "designv11_22_0000822_978-1-4939-0676-5_15-Figure15.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000822_978-1-4939-0676-5_15-Figure15.12-1.png", "caption": "Fig. 15.12 Schematic view of a nanopore electrode", "texts": [ " Since the current varies slightly with the thickness of the ring, it depends almost on the radii rather than on the electrode area. In the recessed electrode, the active surface is located at the bottom of a hole. Simple geometry of the electrode is a disk as is shown in Fig. 15.11. These electrodes can be fabricated by etching the metal wires. The steady-state current for such geometry is expressed by equation T2,15. 32 A geometry that is somewhat similar to a recessed ME is that of the nanopore electrode (Fig. 15.12). This electrode geometry is characterized by the small pore orifice, whose radius, a, can be varied between 5 nm and 1 \u03bcm; the pore depth, d; and the half-cone angle \u03d5. An approximate analytical expression for the steadystate limiting current is given by equation T2,16 in Table 15.2.33 The interest for microelectrodes in which the insulating shields is of thickness comparable to the electrode radius is largely driven by the use of microelectrodes as tips in scanning electrochemical microscopy (SECM)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003537_10775463211013922-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003537_10775463211013922-Figure2-1.png", "caption": "Figure 2. Schematic representation of elastohydrodynamic lubrication contact in mixed lubrication regime.", "texts": [ " In the mixed lubrication regime, the load in rough EHL contact is shared between the lubricating film and asperities. Therefore, it is required to satisfy an additional equation that relates the asperity pressure to the surfaces\u2019 separation. At any point of the contact because of load sharing between the lubricant and the asperities, the total pressure is the sum of the hydrodynamic and asperity pressures p \u00bc ph \u00fe pa (1) where p, ph, and pa are total, hydrodynamic, and asperity pressures, respectively. A schematic representation of EHL contact is shown in Figure 2; it demonstrates the contribution of asperities contact and hydrodynamic forces in load sharing. Greenwood andWilliamson\u2019s model (Greenwood, 1966) is used to obtain surface asperity contact pressure. In this model, the equivalent surface roughness parameters such as average radius of asperities, the standard deviation of asperity heights, and density of asperities are used to define the asperity pressure as pa\u00f0x\u00de\u00bc 2 3 ffiffiffiffiffi 2\u03c0 p n\u03b2\u03c3s ffiffiffiffi \u03c3s \u03b2 r E0 Z \u221e \u00f0h\u00f0x\u00de=\u03c3s\u00de s h\u00f0x\u00de \u03c3s \u00f03=2\u00de e \u00f0s2=2\u00deds (2) where s is the normalized height of asperities by \u03c3s and measured from the mean line of the contact surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001516_iros.2015.7353871-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001516_iros.2015.7353871-Figure2-1.png", "caption": "Fig. 2: States of contact for Peg-in-tube", "texts": [ " The peg maintains a continuous contact with the tube during the 978-1-4799-9994-1/15/$31.00 \u00a92015 IEEE 3538 search stage. Thus, the peg moves in a cone with a halfangle of the vertex equals to the amount of the tilt angle \u03b8, as shown in Fig. 1. This section will investigate the geometrical aspects of such contact cases and will extract the required peg center depth cz from the tube\u2019s top surface, as shown in Fig. 1. A tilted peg when lies outside the hole with an offset (cx, cy, cz), can make contact with the tube in four different states. They are shown in Fig. 2 (a \u2013 d). In each of these figures, a rim-rim contact would mean the inner or the outer edge of the tube and the bottom edge of the peg is in contact. There are two other states in which the peg can make contact with the hole. Firstly, when the curved surface of the peg comes in contact with the tube\u2019s inner rim, as shown in Fig. 2(e), and secondly, when the peg\u2019s bottom cap comes in contact with the tube\u2019s inner curved surface, as shown in Fig. 2(f). Latter two cases do not arise during rotation of the tilted peg, which was done during search procedure adapted here in this paper. Hence it is not analyzed anywhere in this work. It may be noticed that the peg-in-tube process has two additional contact cases compared to a peg-in-hole process [9]. They are 1) when the outer rim of the tube and rim of the peg\u2019s bottom cap comes in contact, as shown in Fig. 2(c); 2) when the peg\u2019s bottom cap face comes in contact with the tube\u2019s outer edge as shown in Fig. 2(d). In this section, we define the peg and the tube surfaces, and their edges before we proceed to establish the conditions of peg lying inside, outside or on the tube. The bottom of the peg was defined by a parametric equation for a 3- Dimensional circle p. This is given by p = c+ ur cos\u03b2 + vr sin\u03b2 (1) where u and v are the unit vectors that lie on the peg\u2019s bottom face with origin at the face center and perpendicular to the vector n \u2261 u\u00d7 v, as shown in Fig. 3. The radius of the peg is r, and c is the position vector of the peg\u2019s bottom center", " The point of intersection of the projected peg\u2019s base and the tube Pi = (xi, yi, 0) was obtained by solving E of (4) and (6). The potential point of contact is the Pi which is nearest to Pl. The possible location of Pi for a sample state of the peg is shown in Fig. 3. The cases discussed here completely defines the different stages of peg and tube contact encountered during the search procedure and the depth cz was extracted. 1) Peg lies inside the tube: The projection of lowermost point of the peg Pl on the tube\u2019s top surface lies within the tube\u2019s inner circle, i.e., the hole, as shown in Fig. 2(a). This is one point contact. However, two point contact is also possible which is not shown. One can check this using x2+ y2\u2212R2 1 < 0 for the point Pl. The point of intersection of the projected ellipse and tube\u2019s inner rim Pi that corresponds to the point of contact (the one which is nearer to Pl) satisfies the equation of the plane (3). Thus the depth cz may be calculated from (3) as cz = [nx(xi \u2212 cx) + ny(yi \u2212 cy)]/nz (8) 2) Peg lies on the tube: The lowermost point lies between the inner and outer circular rims of the tube, as in Fig. 2(b). Thus, Pl makes x2 + y2 \u2212 R2 1 \u2265 0 and x2 + y2 \u2212 R2 2 \u2264 0, and the depth cz from Fig. 3 may be given by cz = r sin \u03b8 (9) 3) Peg lies outside the tube: The lowermost point lies outside the outer circular rim of the tube. Thus, Pl makes x2 + y2 \u2212 R2 2 > 0. In this case the peg may have a rimrim contact case, as shown in Fig. 2(c) or rim-face contact case, as shown in Fig. 2(d). In the case when the peg\u2019s cap comes in contact with the outer tube rim, the line joining the center of the tube O to the point of contact Pt = (xt, yt, 0) becomes parallel to the projected normal of the cap\u2019s plane. This gives rise to the simultaneous conditions xt/yt = nx/ny and x2t + y2t = R2 2 (10) Solving which, we get the point Pt = (xt, yt, 0) as xt = sign(nx) \u221a R2 2/(1 + (ny/nx)2) yt = sign(ny)(ny/nx)xt (11) For, nx = 0, xt = 0. Hence, the two sub-cases may be further investigated as \u2022 Rim-rim contact: Pt lies outside the ellipse and makes numerical value of E > 0, as shown in Fig. 2(c). Thus, Pi i.e., the point of intersection of E and tube\u2019s outer circle, lies on the peg\u2019s bottom rim and it satisfies the cap\u2019s plane equation (3). Hence, from (3) cz = [nx(xi \u2212 cx) + ny(yi \u2212 cy)]/nz (12) \u2022 Rim-face contact: Pt lies inside the ellipse E, i.e., the projected bottom cap of the peg, as shown in Fig. 2(d). Thus for Pt lying on the peg\u2019s cap plane, Pt makes E < 0. On substituting Pt = [xt yt 0] in (3) we get cz = [nx(xt \u2212 cx) + ny(yt \u2212 cy)]/nz (13) 4) Curved surface contact: This happens in case of large tilt angle and small offset as in Fig. 2(e). In our algorithm we did not allow such contact and hence it is not discussed any further. The algorithm makes use of the fact that, when a tilted peg attains a two point contact during rotation, the projection of the peg axis on the tube\u2019s top surface represents the direction of the hole and the peg center reaches the minimum depth cz . Thus, the method of finding the hole direction involves rotating the tilted peg about the axis which is perpendicular to the tube\u2019s top plane and passes through the peg center" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002673_j.matpr.2020.07.229-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002673_j.matpr.2020.07.229-Figure4-1.png", "caption": "Fig. 4. Contour plot of 2nd Mode shape (Lateral mode) of cantilever beam.", "texts": [], "surrounding_texts": [ "While conducting the model analysis of the composite cantilever beam, convergence criteria was achieved after several iterations. Using an iterative method, we increase the number of elements by decreasing the size of elements. To determine the suitable mesh size which produces a precise result, the mesh size was decreased until the natural frequencies for a mode shape converged. For the modal analysis of the beam, the number of nodes obtained were 11,028 and the number of elements were 10,983 for an element size of 3 mm. a meshed model of the beam is shown in Fig. 2." ] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure5.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0000617_b978-0-12-814866-2.00007-5-Figure7.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000617_b978-0-12-814866-2.00007-5-Figure7.12-1.png", "caption": "FIG. 7.12", "texts": [ " These sources produce continuous radiation approximating that of a black body radiation. As glass and fused silica are opaque at wavelengths greater than 2.5 mm, the glowing source must not be in a glass bulb or in a casing that is made of these substances. The heated material is either silicon carbide (SiC, called Globar) or is a mixture of rare-earth oxides (producing a device known as a Nernst globar). For a globar heated at 1300e1500 K, useable light is provided at wavelengths of 0.4e20 mm. (I) The Nernst Glower: The general construction of the Nernst glower is shown in Fig. 7.12A. This design includes the semiconducting material made of rare-earth oxide formed into a cylinder having diam. of 1e3 mm and length of 20e50 mm. Platinum leads are sealed to the ends of the cylinder to permit electrical connections. As current passes through the device, its temperature is increased to between 1200 and 2200 K. A heating source and a reflector is included to help pass the radiation coming out from the light source in the desired direction. Because the Nernst glower has a large negative temperature coefficient of electrical resistance (resistanceY as temperature[) thus circuit is current limited. Therefore, it must be heated externally to a dull red hot before the current is large enough to maintain the desired temperature. This device produces adequate amounts of light at wavelengths ranging from 1 to 40 mm. Fig. 7.12B shows a typical spectral distribution, and shows the light produced by such a device closely matches what would be expected for blackbody radiation. (II) Silicon carbide Globar Introduction: Globar is a thermal light source that emits radiation in near IR region of the electromagnetic spectrum. Globar is a thermal radiator, meaning a solid which is heated to very high temperatures to produce light radiation. The name \u2018globar\u2019 is simply a combination of glow and bar, the heating element in the form of a bar is heated until it glows to emit radiation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001277_acc.2014.6858589-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001277_acc.2014.6858589-Figure4-1.png", "caption": "Fig. 4. Inverted Pendulum Controlled by DC Motor", "texts": [ " Here, we see that the original plant is subjected to process noise w(t) and measurement noise v(t). The entire algorithm of combined estimation and control leading to nonlinear tracking problem is shown in Table I where we introduced the command r(t) for a more general treatment [15]. V. ILLUSTRATIVE EXAMPLE For numerical simulation and analysis, the developed estimation and optimal tracking technique is implemented for noise cancellation for inverted pendulum controlled by DC motor, as shown in Fig. 4 The dynamic equations for system under concern are V (t) = L di(t) dt +Ri(t)+ kb d\u03b8 (t) dt , (43) ml2 d2\u03b8 (t) dt2 = \u2212mglsin(\u03b8 (t))\u2212 kmi(t), (44) where, V is the control voltage, L is the motor inductance,i is the current through the motor winding, R the motor winding resistance,kb the motor\u2019s back electro magnetic force constant, \u03b8 the angle of pendulum, m the mass of pendulum, l the length of rod, g the gravitational constant, and km the damping (friction) constant. The system nonlinear state equations can be written in the state dependent form x\u03071 x\u03072 x\u03073 = 0 1 0 (g/l)sin(x1) x1 0 km ml2 0 \u2212 kb L R L x1 x2 x3 + 0 0 1 L u, (45) where: \u03b8 = x1 , \u03b8\u0307 = x2 , i = x3, V = u" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002438_j.actaastro.2020.04.063-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002438_j.actaastro.2020.04.063-Figure1-1.png", "caption": "Fig. 1. Mission concept of constructing a large-scale antenna by ISA.", "texts": [ " However, if size of the structure exceeds the order of 100 m, usually it needs be designed as a multi-module structure and constructed through in-space assembly (ISA). Over the past decades, with remarkable advances in the development of assembly robotic and autonomous operations, the ISA missions such as modular space telescopes [4] and space solar power stations [5], have received extensive attention [6,7]. The large antenna construction using ISA is of particular interest in this research. As shown in Fig. 1, a large-scale antenna could be integrated in-orbit with modules by using assembly robot. The size of entire antenna could be as large as 500 m \u00d7 3 m while each module is designed at 5 m \u00d7 3 m in size. As shown in Fig. 2, the assembly robot consisted of two major parts: a rigid central service body and an adhesive device [8]. The adhesive device was designed according to the principle of gecko claw bionics [9]. It can catch a module and transport to the specified position for integration, and then comes back to catch another module" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003619_tie.2021.3080207-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003619_tie.2021.3080207-Figure1-1.png", "caption": "Fig. 1. 10-pole/12-slot DTPSPM machines. (a) Symmetrical winding configuration. (b) Asymmetric winding configuration. (c) Phasor diagram of phase back-EMF for symmetrical winding configuration. (d) Phasor diagram of phase back-EMF for asymmetric winding configuration.", "texts": [ "6% larger than the smallest value under different start rotor positions, which can result in much difference of demagnetization. In [7], it was also revealed that the variation of the d-axis 3PSC peak current repeats for every multiple of \u03c0/3 in an electrical period and the d-axis peak current occurs when 3PSC happens at the rotor positions k\u03c0/3 (k=0, 1, 2\u2026). However, the reasons for such rules were not explained and whether the conclusions are applicable to other PM machines was not pointed out. This article is dedicated to fill in this gap. Since symmetrical DTPSPM machines, such as the one shown in Fig. 1(a), have similar 3PSC response with single three-phase PM (STPPM) machines [1], the answer of above problems will be applicable to both machine types. Moreover, some DTPSPM machines have asymmetric mutual inductances within one winding group, such as the one shown in Fig. 1(b). In such asymmetric DTPSPM machines, the conclusions in [7] can be inapplicable. In addition, little literature was found regarding the influence of start rotor position on the q-axis 3PSC peak current. This article aims at guiding which rotor position should be selected to start 3PSC when checking the reliabilities at the design stage of DTPSPM machines. For example, as aforementioned, irreversible demagnetization is usually checked under 3PSC currents in DTPSPM machines, especially those for safety-critical applications", " INFLUENCE OF START ROTOR POSITION ON 3PSC IN ASYMMETRIC DTPSPM MACHINES A. Asymmetric Mutual Inductances Dual three-phase PM machines can usually be seen as STPPM machines allocating their windings into two winding groups. When windings of STPPM machines are spatially divided into even number of modules, which consist of the same number of unit machine or half unit machine, and every other module is designed into one winding group, the asymmetry of mutual inductances occurs within one winding group. As shown in Fig. 1(b), the 10-pole/12-slot machine is divided into two modules, which contains half unit machines. Since the coils A1 and C1 at the two ends of each module is spatially discontinuous, the mutual inductance between the two coils is almost 0H and thus the mutual inductance Mac is smaller than Mab and Mbc within single winding group. Therefore, the inductances within single winding group are 1 1 1 1 1 1 1 1 0 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 0 1 1 1 A A A B A C B A B B B C C A C B C C L M M L M M M M L M M L M M M L M M M L \u2212 \u2212 = \u2212 \u2212 \u2212 \u2212 (15) in the phase coordinate system and ( ) ( ) ( ) ( ) 1 1 1 1 0 1 1 1 1 1 1 0 00 0 0 1 0 0 13 sin 2 2 / 6 cos 2 2 / 62 cos 2 2 / 6 sin 2 2 / 63 d d d q d q q q L M M L M M L t tM t t = + \u2212 + + + + \u2212 + + \u2212 + + (16) in the dual dq coordinate system, where L1 and M1 are the Authorized licensed use limited to: Carleton University", " In addition, the absolute difference between the largest and smallest 3PSC peak currents increases with asymmetry level, as shown in Fig. 11 (b). IV. FE AND EXPERIMENTAL VALIDATION A. FE Validation In order to verify the above analysis, three-dimensional FE (3D-FE) simulation is performed on a 10-pole/12-slot DTPSPM machine. The 3D-FE model is shown in Fig. 12. The parameters are listed in TABLE V and the B-H curve of laminated steel sheets is shown in Fig. 13. Firstly, the influence of start rotor position on 3PSC peak currents for the symmetrical winding configuration shown in Fig. 1(a) is simulated. Both linear and nonlinear core materials are considered. It is assumed that the initial current of faulty winding and the healthy winding current are 0A whilst the rotor speed is 5000rpm during 3PSC. Firstly, the spectra of open-circuit flux linkage of single winding group are simulated, as shown in Fig. 14. When core saturation is not considered, k5 equals to 1 and k7 equals to -1. Based on (10) and (13), the largest d- and q-axis peak currents occur around the start rotor position 0 and 13\u03c0/180, respectively", "org/publications_standards/publications/rights/index.html for more information. positions leading to the largest d- and q-axis 3PSC peak currents are about \u03c0/6 and 5\u03c0/24, respectively. It is noted that although the angle between two winding groups for this symmetrical winding configuration is 30 electrical degrees, the conclusions are also applicable to the other angle displacements. Based on the same FE model, the influence of start rotor position on 3PSC peak currents for asymmetric winding configuration shown in Fig. 1(b) is simulated with the same assumptions. As shown in Fig. 17, when core saturation is neglected, the start rotor positions leading to the largest 3PSC peak currents match well with the analytical results. Therefore, the analytical model is validated. In addition, under such rotor speed, it can be seen the d-axis 3PSC peak current is larger than the q-axis peak current at each start rotor position. The d-axis peak current also has larger difference between the largest and smallest values than the q-axis peak current" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001977_j.ijfatigue.2016.09.006-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001977_j.ijfatigue.2016.09.006-Figure2-1.png", "caption": "Fig. 2. Characteristic geometrical properties of a fillet welded joint, where q is the toe radius and h is the weld toe blend angle.", "texts": [ "2% plastic strain (MPa) Rm ultimate tensile strength (MPa) t weld plate thickness (mm) Zmin minimal reduction in area after tensile test (%) a, b exponents in the Kf (i) (N) function (\u2013) m Poisson\u2019s ratio (\u2013) h weld face angle (deg) q notch root radius (mm) ra stress amplitude (MPa) ra (i) (N) stress amplitude in the S-N curve of parent material under load i, corresponding to a given number of cycles N (MPa) ra (w,s) (i) (N) nominal stress amplitude in the S-N curve of welded test pieces under load i (with geometric/structural notch - respectively), corresponding to a given number of cycles N (MPa) rmax stress at the notch root (MPa) rn nominal stress (MPa) For welded joints a typical assumption is that the Kt factor depends on weld notch root radius, q, and on the weld toe blend angle h (Fig. 2). As can be seen, such an approach does not involve the effect of material heterogeneities on stress and strain tensor [22]. The issue of notch factor exists also in the fatigue life determination procedures, where the term of fatigue notch factor Kf is known. Such a factor describes differences between S-N curves corresponding to parent material and to the considered welded element (Fig. 3). These differences result from the differences in stress fields; therefore the approach to Kt determination indirectly define also the approach to determination of fatigue notch factor, Kf" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001850_amr.1137.61-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001850_amr.1137.61-Figure7-1.png", "caption": "Fig. 7 (a) Different tool tips (right angle and chamfered corner), (b) Diamond drills for rotary USM", "texts": [ " Coolant pumped through the core of the drill washes away the swarf, prevents jamming of the drill, and keeps it cool [58]. High material removal rate is achieved during Rotary USM as compared to conventional one [1]. The various process variables in rotary USM are rotational speed, vibration amplitude and frequency, diamond type, size and concentration, bond type, coolant type, and pressure etc. The various machining characteristics are material removal rate, cutting force and surface roughness etc. Fig. 7 and 8 shows the rotary ultrasonic machine tool. Advanced Materials Research Vol. 1137 69 Churi et al. [58] studied the effects of different process parameters (ultrasonic power, spindle speed and feed rate) on the machining characteristics of titanium alloy in rotary USM. The results reported that feed rate has significant effect on cutting force, MRR and SR. Increase in power rating reported with decrease in surface roughness. Lee et al. [59] optimized the process parameters in rotary ultrasonic machining of glass lens using taguchi method" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001871_978-3-319-32098-4_34-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001871_978-3-319-32098-4_34-Figure4-1.png", "caption": "Fig. 4 A FEA model of the Ti64 lattice structure", "texts": [ " For each structure, the simulation approach, called the continuum elements model, was carried out. To reduce the computation time, one-quarter of the symmetrical model of each structure was built. The multi-units model could thus reveal the interaction between the adjacent units under compressive loading conditions. Two rigid surfaces called top and bottom, respectively, were modelled in the part module, and they were used to simulate the crosshead and substrate in the compression tests. Rigid and non-friction contact conditions were set between the rigid surfaces and the structure. As Fig. 4 shows, two reference points (RPs) were placed at the centre of the rigid surfaces. The one on the top surface applied the displacement boundary conditions to the models, and during the simulation, the displacement of the top surface was recorded. The reaction force applied in the RP on the bottom surface was also recorded. A dynamic explicit analysis step was established for each FEA model after the default initial step. In this step, called \u2018press\u2019, the upper rigid surface shown in Fig. 4 moved downwards and compressed the structure, while the bottom rigid surface was fixed during the whole process. Moreover, the material models were assumed to be isotropic, and the von Mises yield criterion was used in the FEA analysis. The material parameters were obtained via the above tensile tests, and the resultant nominal stress\u2013strain curve is shown in Fig. 5a. The true stress\u2013strain data converted from the nominal stress\u2013strain data were used as the material properties in each FEA process" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003596_j.jbiomech.2021.110527-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003596_j.jbiomech.2021.110527-Figure2-1.png", "caption": "Fig. 2. Diagram of the two SLIP models (a) the reference model at touchdown (i.e., h = p-a reference model has two degrees of freedom. One degree of freedom is the primary mass angle, h, of the spring with respect to the vertical. The attack angle, a, is the angle of the three degrees of freedom. The degrees of freedom include the primary mass radial positio to the vertical (not depicted).", "texts": [ " This addition to the SLIP model was designed to mimic the wobbling soft tissues of humans. The revised SLIP model was compared with the regular SLIP model (Fig. 1) for its ability to self-stabilize, and its robustness to perturbations. Two SLIP models were utilized to access the effects of early stance phase energy dissipation. A glossary of model variables and parameters can be found in Table 1. The reference model (RM) was the same model described in Seyfarth et al. (2002). The equations of motion for the RM (Fig. 2) during the stance phase are, \u20ach \u00bc g sin h\u00f0 \u00de 2_rleg _h rleg \u00f02\u00de The experimental model (EM) included an additional model component, a mass-spring-damper mechanism designed to represent the wobbling masses (Fig. 1b). The equations of motion for the EM are, \u20acrleg \u00bc gm1 cos h\u00f0 \u00de kleg\u00f0rleg lleg\u00de \u00fe k2\u00f0r2 rleg\u00de \u00fem1rleg _h2 m1 \u00f03\u00de \u20acr2 \u00bc gm2 cos h\u00f0 \u00de k2 r2 rleg \u00fe c2 _r2 \u00fem2r2 _h2 m2 \u00f04\u00de where, c2 \u00bc c2 tanh b\u00f0rleg lleg\u00de ; _r2 < 0 0; _r2 0 . \u20ach \u00bc g sin h\u00f0 \u00de\u00f0m1rleg \u00fem2r2\u00de 2\u00f0m1rleg _rleg \u00fem2r2 _r2\u00de _h m1rleg2 \u00fem2r22 \u00f05\u00de The tanh term is necessary to enforce zero damping at touchdown and takeoff" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000154_j.wear.2018.12.088-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000154_j.wear.2018.12.088-Figure16-1.png", "caption": "Fig. 16. Weight loss vs running distance of adhesive wear tests.", "texts": [], "surrounding_texts": [ "The dry sliding gear wear test was performed as the worst scenario case baseline for worn surface and wear debris analysis and then compared to other wear modes. In general, the worn surface in comparison to an unworn surface had a severe adhesive wear pattern as shown in Fig. 24(a) and (b). In conjunction with that, severe adhesive wear particles, as shown in Fig. 25 were observed and confirmed that normal rolling and sliding occurred during the test performed under a similar condition to other tests. The wear debris was firstly extracted from its lubricating oil sample, this was achieved by filtration using membrane patch filter paper with 0.45 \u00b5m pore size. Fig. 20 shows a typical worn out surface of 10,000m running distance without lubricant supply. The worn surface exhibits failure wear modes such as scuffing and scoring. In conjunction to the worn surface, typical severe adhesive wear debris was observed as shown in Fig. 25. This is evidence that direct metal to metal contact was the major wear mechanism whereby the asperities of the two opposing gear surfaces were plastically deformed and weld together leading to transfer from one surface to the other." ] }, { "image_filename": "designv11_22_0003398_s00466-020-01960-9-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003398_s00466-020-01960-9-Figure5-1.png", "caption": "Fig. 5 A comparison of stress and temperature contour results for a laser beam scan on a bare plate of 304L stainless steel for a the part-scale process model and b the high-fidelity multiphysics workflow", "texts": [ " The computational time and resources required for the highfidelity multiphysics workflow allow the incorporation of the effects of melt flow, gas dynamics, surface tension and vapor recoil pressure. However, the thermal/fluid simulations can vary significantly based on how complex the melt flow becomes. The high-fidelity multiphysics workflow also includes an interfacial conforming mesh allowing the surface topology to evolve throughout the simulation providing amore realistic resultingmorphology. Figure 5 shows a comparison of the previous part-scale workflow and the proposed high-fidelity multiphysics workflow for the bare plate laser scan example problem. As shown in Fig. 5, the high-fidelity multiphysics workflow results show higher peak stress and a spatially varying stress field in addition to higher temperatures compared to the part-scale model. Although, predicted stresses show similar trends between both workflows, the variance and higher peak stresses as shown in Fig. 5 for the high-fidelity multiphysics workflow demonstrate the effects of additional physical considerations captured in the high-fidelity multiphysics workflow. Comparisons of the resulting cross sections and temperature distributions for part-scale and high-fidelity multiphysics workflows are compared with an experimental cross section of the same bare plate laser scan in Fig. 6. Temperatures shown are maximum temperatures over the length of the simulation so that melt depths can be determined and the melt depth prediction is from a suite of thermal uncertainty quantification runs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001091_05698196108972425-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001091_05698196108972425-Figure1-1.png", "caption": "FIG. 1. Geometry of displaced center two-lobe bearings.", "texts": [ " They may be named, respectively, cylindrical , displaced cylindrical , elliptical , and displaced-elliptical two-lobe bearings, and are illustrated diagrammatically in Fig. 2. [3] m + \u00a3 cosrp --- - [5] \u00a31 COS>< >>: \u00f014\u00de where Ib can be obtained by equation (15) while ignoring its variety along the moulded line direction Ib \u00bc D2 b D2 f 4NLst Is \u00f015\u00de Similar to the finger beam field, the RMV of the finger foot field is selected as shown in Figure 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003433_s12555-019-0962-z-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003433_s12555-019-0962-z-Figure3-1.png", "caption": "Fig. 3. A quadrotor configuration.", "texts": [ " To pursue increasingly reliable performance, it is necessary to reduce \u03b1 and set it away from critical value. Reconsidering the limit \u03b1 < 1+ ln\u039b/(ln\u03b4 \u2212 ln(1+\u03b4 )) and \u039b = eFT , we can find that once sampling period T decreases, \u039b will decrease correspondently. In practice, if a big package loss rate \u03b1 is evitable, choosing a smaller sampling period may be a good option. 4.1. Adaptive quantized control for UAV with Lipschitz logarithmic quantizer The quadrotor UAV, whose configuration is depicted in Fig. 3 [24], is driven by four propellers respectively in- stalled at the end of each arm, and can track time-varying translational and rotational command signals by changing voltage inputs of DC motors. To verify the effectiveness of the adaptive control with quantized input processed by Lipschitz logarithmic quantizer, we introduce a quadrotor UAV, a sophisticated multichannel control target with six state variables p= [x,y,z]T , \u0398 = [\u0393\u03b8 ,\u0393\u03c6 ,\u0393\u03c8 ] and four practical operating control signals [U1,U2,U3,U4] = [Up,U\u03b8 ,U\u03c8 ,U\u03c6 ]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002725_tmag.2020.3021644-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002725_tmag.2020.3021644-Figure2-1.png", "caption": "Fig. 2. Topology of the proposed DFM-CMG.", "texts": [ " Structurally, each of the CMGs meets [1, 15] LHm PPZ (1) where PH is the PPN of HSR PMs, PL is the PPN of LSR PMs, and Zm is the number of modulator segments. The air-gaps are marked in Fig. 1. The magnetic field in airgap II of CMG I contains the PH th and PL th order harmonics. If the HSR of CMG II is replaced with the HSR and modulator of CMG I, the PH th order harmonic in air-gap II can work as a HSR for CMG II, and the PL th order harmonic in air-gap II can couple with LSR PMs of CMG II. Hence, the DFM-CMG shown in Fig. 2 can be regarded as a combination of CMGs I and II. Due to the integration of two CMGs, the DFM-CMG structure is more complex. It consists of a HSR, an outer modulator (OM), a LSR, and an inner modulator (IM). The number of OM segments, Zm, is the same as the number of IM teeth. It can be configured by using (1). The relative angular displacement between the two modulators Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 20,2020 at 21:04:18 UTC from IEEE Xplore" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001083_1.4027926-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001083_1.4027926-Figure4-1.png", "caption": "Fig. 4 Impulse experiment with aero-engine rotor experimental rig", "texts": [], "surrounding_texts": [ "4.1 Faults Produced on Ball Bearings. To study the casing responses caused by ball bearing faults, in this study, faults are produced on ball bearings using wire-electrode cutting technology. These faults are produced on the inner ring, outer ring, and ball of the compressor experimental rig\u2019s 6214 ball bearing and the aero-engine experimental rig\u2019s 6205 ball bearing, which are shown in Figs.7 and 8. The ball bearing dimensions are listed in Table 1. The ball bearing faults\u2019 characteristics frequencies can be computed as follows: (1) Outer race foc \u00bc Z 2 1 d D cos a fR (1) (2) Inner race fic \u00bc Z 2 1\u00fe d D cos a fR (2) (3) Rolling ball fbc \u00bc Z 2 D d 1 d D cos a 2 \" # fR (3) (4) Cage fc \u00bc 1 2 1 d D cos a fR (4) 4.2 Faults Experiments of Ball Bearings. The bearing faults experiments are carried out. The test-site photo of the compression rotor rig is shown in Fig. 9, and the measurement points\u2019 explanation is listed in Table 2; the test-site photo of the aeroengine rotor rig is shown in Fig. 10, and the measurement points\u2019 explanation is listed in Table 3. Vibration signals are collected by means of the USB9234 data acquisition card of the NI Company, the 4805 type ICP acceleration sensors of B&K Company are used to pick up the acceleration signals, and the eddy current sensors are used to measure the rotating speeds. The sampling frequency is 10.24 kHz. 4.3 Wavelet Envelope Analysis for Ball Bearing Fault 4.3.1 Basic Principles of Wavelet Envelope Analysis. In this study, a signal analysis to determine ball bearing faults is carried out by means of a wavelet envelope spectrum analysis. Chen [2] provides a reference for the detailed process of this algorithm. The essence of ball bearing fault diagnosis based on the wavelet packet is to take advantage of its bandpass filter characteristics, and to decompose the signals using appropriate wavelet functions so as to obtain an appropriate resonance frequency band. Then, by means of envelope demodulation, low frequency envelope signals that only contain the fault characteristic information are obtained. Its spectrum is the wavelet envelope spectrum, in which the fault characteristic frequencies of the ball bearings can be found out. 4.3.2 Wavelet Envelope Spectrum Analysis of Ball Bearing Fault Signals 4.3.2.1 Experiment Analysis Based on Compressor Rotor Experimental Rig (1) Feature extraction for inner ring faults Figures 11\u201314, respectively denote the time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope spectra of the bearing house\u2019s response and the casing\u2019s response to inner ring faults in the 6124 ball bearing. The experimental rotating speed is 1793 rpm (29.88 Hz). The number of balls is 10, the other ball bearing\u2019s parameters are listed in Table 1, and the inner ring characteristic frequency can be calculated by formula (2) as fic\u00bc 5.8974 29.88\u00bc 176.24 Hz. From Fig. 11, the vibration acceleration response value of the casing is larger than that of the bearing housing, which shows that the impulse vibration of the bearing fault is effectively transmitted to the casing. This result is in full agreement with the experimental result for the impulse response. From Fig. 12, it can be found out that there are many resonance peaks in the frequency spectrum, and the signals are very weak in the low frequency segments, the signal\u2019s frequency spectrum from 0 Hz to 500 Hz are shown in Fig. 13, from which, the characteristic frequency of the inner ring fault and its modulation frequency fr can be Journal of Dynamic Systems, Measurement, and Control NOVEMBER 2014, Vol. 136 / 061009-3 Downloaded From: http://dynamicsystems.asmedigitalcollection.asme.org/ on 01/29/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use basically seen. The wavelet envelope spectrum is shown in Fig. 14, by comparison with Fig. 11, the characteristic frequency of the inner ring fault and the modulation frequency fr are more distinctly shown in Fig. 14. Obviously, the inner ring fault features can be all extracted from the acceleration signals of the bearing house and the casing by means of the wavelet envelope spectrum analysis and frequency spectrum. (2) Feature extraction of outer ring faults The time domain waveforms, frequency spectrum (0\u20135000 Hz), frequency spectrum (0\u2013500 Hz), and wavelet envelope spectrum of the bearing house responses and casing responses of the 6124 ball bearing outer ring faults are respectively shown in Figs. 15\u201318. The experimental rotating speed is 1826 rpm (30.43 Hz). The number of balls is 11, the other ball bearing\u2019s parameters are listed in Table 1, and the outer ring characteristic frequency can be calculated by formula (1) as foc\u00bc 4.5879 30.43\u00bc 140 Hz. From Fig. 15, the vibration acceleration response value of the casing is larger than that of the bearing housing, which shows that the impulse vibration of the bearing fault is effectively transmitted to the casing. This result is in full agreement with the experimental result for the impulse response. From Fig. 16, it can be found out that there are many resonance peaks in the frequency spectrum, and the Table 1 Ball bearing dimensions (in mm) Type Diameter of inner ring Diameter of outer ring Thick Diameter of ball Pitch diameter" ] }, { "image_filename": "designv11_22_0002596_1350650120942009-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002596_1350650120942009-Figure1-1.png", "caption": "Figure 1. Slider crank mechanism dynamic model and its impacting force on the crankpin bearing. P is the combustion gas force impacted on the piston peak; L and R are the connecting rod length and rotation radius of the crankshaft; and ! is the angular velocity of the crankshaft.", "texts": [ " The dimensions of the designed textures are then optimized based on the multi-objective genetic algorithm (MOGA) to enhance the LE-FPL. This study\u2019s innovation is the design and optimization of the lubrication textures on the bearing surface to further improve the LE-FPL. Lubrication model of surface pairs In the working process, the crankpin bearing including a crankpin surface moves on a bearing surface. Under the impact of the dynamic force of the slidercrank mechanism, the LE-FPL of crankpin bearing can be influenced. Thus, the dynamic model in Figure 1 is used to calculate the slider-crank mechanism\u2019s impacting force W0 on the crankpin bearing. Based on the reference data of the combustion gas pressure in the engine at 2000 r/min and the motion equation of the slider-crank mechanism,6 the result of the impacting force of W0 is calculated and plotted in Figure 2. Under the impact of W0 changed in both direction and intensity, the load-bearing capacity of crankpin bearing (Wb) generated by the load-bearing capacity of the oil film pressure (Wof) and asperity contact (Wac) in the mixed lubrication region must be equal toW0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001170_0278364914551773-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001170_0278364914551773-Figure3-1.png", "caption": "Fig. 3. Frames for WMR kinematics. The four frames necessary for the relation of wheel rotation rates to vehicle speed and angular velocity.", "texts": [ " Once the contact point is defined and its velocity is known, it is necessary to invoke knowledge of wheel slip over the terrain in order to determine the associated rotations of the wheels. Let the actual contact point velocity v *w c a be composed of an ideal no-slip component v *w c i derived from wheel rotation rate, plus a slip velocity dv *w c v *w c a = v *w c i + dv *w c \u00f054\u00de Once the slip velocity is known, the actual velocity can be used in the wheel equation for estimation and the ideal velocity can be used for control. Consider Figure 3. For illustration purposes, we will allow a single degree of articulation between the vehicle and contact point frames. Let us define the following frames of reference: w, world, fixed to the environment; v, vehicle (body), fixed to some point on the vehicle whose motion is of interest; s, steering, positioned at the hip/steering joint, moves with the boom (if any) to the wheel; c, contact point (wheel), moves with the contact point, has the same orientation as the steering frame. Each frame has a default associated Cartesian coordinate system, based on the axes shown in the figure" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001651_s00170-016-8545-0-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001651_s00170-016-8545-0-Figure6-1.png", "caption": "Fig. 6 EDM machine tool", "texts": [ " In the visual coordinate system, direction vector component p represents the weight along the main optical axis. When p = 0, it means that the two straight lines lb and ls are vertical to the main optical axis of camera. Thus, the camera should be kept at the posture which makes main optical axis vertical to end face of the blisk blank. Finally, the camera is placed on a rotary table which has two rotational freedoms to adjust its main optical axis vertical to end face of blank. In this way, operation can be simplified and it also makes use of the high accuracy of the EDM machine. Figure 6 shows the blisk placing on the machine tool, which has six axes. After the pictures of the blisk blank and the spindle are taken, there will be several steps to get useful information from them. The first step is graying. Since the features in the pictures have nothing to do with color. It will be more convenient to process the pictures after removing the colors. The second step is filtering work to reduce noises. The noises mainly appear as some high light pixel blocks or isolated pixels in the pictures" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001656_tmag.2014.2330812-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001656_tmag.2014.2330812-Figure5-1.png", "caption": "Fig. 5. Prototyped M1 and M2. M1 is a simple spherical NdFeB magnet. M2 is composed of four cylindrical NdFeB magnets.", "texts": [ " Using this procedure, the SPRMM can be precisely manipulated along a complex predefined moving path on a complex 3D thin surface. We performed various experiments to verify the proposed control method. We first constructed an MNS, as shown in Fig. 4 and Table I [8]. The MNS is composed of an x-directional HC, a y-directional USCy, and a z-directional USCz. Two orthogonal (biplane) cameras (xy plane: Cam1, xz plane: Cam2, resolution: 640 \u00d7 480, sampling rate: 10 Hz) were installed to obtain real-time 3D images of the working space inside the MNS. We also prototyped the SPRMM as shown in Fig. 5. M1 is a simple spherical NdFeB, and M2 was constructed by 3D printing technology using ultraviolet curable acrylic plastic material. Four cylindrical NdFeB magnets (n = 4) were periodically inserted in M2 [5]. The magnetization of the NdFeB magnet is 9 55 000 A/m. To make the SPRMM continuously roll along a surface without slip, M1 and M2 were built with the geometric parameters shown Fig. 5 [5], and the outer surface of M2 was painted blue so that it could be effectively tracked by blue color image tracking. We first observed the rolling locomotion of the SPRMM along a predefined spiral rolling path on a horizontal thin surface using open-loop control and the proposed closed-loop control methods, as shown in Fig. 6. Sequential spiral moving paths of the SPRMM were manually set with 150 sequential target points using a graphic interface in the control panel. We set the rotating speed and magnitude of the ERMF as 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001651_s00170-016-8545-0-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001651_s00170-016-8545-0-Figure5-1.png", "caption": "Fig. 5 Projection transformation", "texts": [ " After the blisk is placed on the EDM machine tool, two coordinate systems need to be set up for visual identification. One is the world coordinate system (WCS) which is the same as the EDM machining coordinate system, and the other is the visual coordinate system(VCS). In the VCS, the main optical axis of the camera is set as the N-axis, and the horizontal and vertical of the camera are U-axis and V-axis respectively. The origin of the VCS is the center of camera lens. The projection transformation of the camera is shown in Fig. 5. A reference of theWCS, such as the spindle, and the blisk blank shall be taken into the same picture. Thus, a positional relationship between the blisk and the WCS can be found in the picture. While the coordinator transformation from the WCS to the VCS is established, the real positional relationship is able to be finalized. Then, the blisk could be rotated to the circumferential initial position for EDM machining. The coordinator transformation from the WCS to the VCS is shown in Eq. 1. [ xv yv zv 1 ] = [ xw yw zw 1 ] \u2022 T (1) where (xv, yv, zv) and (xw, yw, zw) are the coordinates values of the same point in the VCS and the WCS respectively, and T is the transformation matrix. T commonly contains both rotation transformation and translation transformation. Homogeneous transformation matrix is used here to describe this conversion process of the coordinate systems. The transformation matrix T is expressed as T = \u23a1 \u23a2 \u23a2 \u23a3 Ux Nx Vx 0 Uy Ny Vz 0 Uz Nz Vz 0 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a6 \u2022 \u23a1 \u23a2 \u23a2 \u23a3 1 0 0 0 0 1 0 0 0 0 1 0 dx dy dz 1 \u23a4 \u23a5 \u23a5 \u23a6 (2) As shown in Fig. 5, the imaging plane of the camera is vertical to the main optical axis. The distance between the imaging plane and the VCS origin is defined as the focal length of the camera. According to perspective projection, the coordinate values of the points in the imaging plane can be gotten by applying projection transformation to VCS as Eq. 3. xi L = xv zv , yi L = yv zv (3) Where (xi, yi) is the point in the projection plane and L is the focal length of camera. From Eqs. 1\u20133, coordinate transformation from the WCS to the VCS could be built" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000709_psc49016.2019.9081455-Figure22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000709_psc49016.2019.9081455-Figure22-1.png", "caption": "Fig. 22. topology of electromagnetic suspension [17].", "texts": [ " It could be shown that in order to increase the powerfactor and efficiency by reducing the leakage flux of LIM, edges of the conductive track can be curved as illustrated in Fig. 20 [15]. Another way of reducing the end effects is by using the cylindrical design which is shown in Fig. 21 [16]. IV. SUSPENSION CONTROL IN MAGLEV TRAINS As electrodynamic suspension maintanes the air gap without needs for a control system, in this chapter, only the electromagnetic suspension is considered. The topology of the electromagnetic suspension is demonstrated in Fig. 22. It consists of a set of suspension soils which depending on their positions on the left, right , front and rear sides of the bogie, could be divided into 4 main groups [17]. Depending on how the aforementioned coil groups are controlled, there are three types of control strategies. In this chapter, these strategies are discussed. A. Single independent control of groups In this strategy, each bogie is considered to have two groups of front and rear coils, each of which is fed through an independent control system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003537_10775463211013922-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003537_10775463211013922-Figure3-1.png", "caption": "Figure 3. Experimental test setup and a block diagram of the measurement and data handling system.", "texts": [ " The remaining unknown parameters of the proposed modified LuGre model, that is, \u03c30 and \u03c31, which define the contact mico-slip behavior, are identified using the system\u2019s measured behavior. In the following, an experimental case study demonstrates the properties of the proposed model, and its advantages over the existing models in the literature are discussed. To validate the presented model of lubricated roller guideways and the capabilities of the proposed modified LuGre model in predicting contacts friction behaviors, an experimental case study was conducted. Figure 3 shows the experimental test setup and a block diagram of themeasurement and data handling system consisting of a linear roller guideway system (INA RUE45-E-H, Schaeffler Group, recirculating linear roller bearing guideway, 2018). In the setup, the carriage was fixed, whereas a shaker drove the rail. Single harmonic excitation forces at low frequencies were applied to the rail. The force and response signals were measured with the sampling rate of 1600 samples per second and transferred to the data acquisition system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000465_0954406219878755-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000465_0954406219878755-Figure2-1.png", "caption": "Figure 2. Discoid conical grinding wheel in a.", "texts": [ " By means of this technique, the multivariable nonlinear equation group to compute the meshing limit line of the worm drive may be resolved smoothly. In the current work, the computing principle of the meshing limit line is well built for the conical worm pair. Based on the preceding theory and the method suggested, the meshing limit line characteristics of the enveloping conical worm pair are investigated in detail. The numerical example study is also performed. Geometric model of enveloping conical worm pair Equation of grinding wheel generating surface As illustrated in Figure 2, a right-handed coordinate system a Oa; ~ia, ~ja, ~ka is rigidly connected to a discoid conical grinding wheel. The vector ~ka of a coincides with the axis of the grinding wheel. The two symbols g and rg denote the half taper angle and the radius of the grinding wheel, respectively. Consequently, the equation for the grinding wheel generating surface, g, can be determined in a by the aid of the sphere vector function,4,5 and the outcome is ~ra a \u00bc u ~ma , g \u00bc u sin g cos ~ia \u00fe u sin g sin ~ja \u00fe u cos g ~ka, u4 0, 04 5 2 \u00f01\u00de in which the symbols u and are the two surface parameters of g. The unit normal vector to g can be obtained as ~n a \u00bc @ ~ra\u00f0 \u00dea @ @ ~ra\u00f0 \u00dea @u @ ~ra\u00f0 \u00dea @ @ ~ra\u00f0 \u00dea @u \u00bc ~na , g \u00bc cos g cos ~ia \u00fe cos g sin ~ja sin g ~ka \u00f02\u00de Obviously, the orientation of ~n a achieved above is from the grinding wheel entity to the space as displayed in Figure 2. Generation of enveloping conical worm As described in Figure 3, a fixed coordinate system o1 O1; ~io1, ~jo1, ~ko1 is affiliated with the worm rough. The origin O1 of the coordinate system o1 locates at the middle point of the thread length of the worm and the unit vector ~ko1 lies along the worm rough axis. A rotating coordinate system 1 O1; ~i1, ~j1, ~k1 is rigidly linked to the worm rough and is utilized to denote its present location. The rotating angle of 1 relative to the coordinate system o1 is the angle \u2019" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002258_s00170-020-04953-3-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002258_s00170-020-04953-3-Figure3-1.png", "caption": "Fig. 3 Synthesis of velocity", "texts": [ " There are two special circumstances: one is when \u03c6 is fixed, xM and yMwill change with the time t, which means that the planeMN1N2 will do translational movement on the fixed plane; the other is when xM and yM are fixed, \u03c6 will change with the time t, indicating that the plane MN1N2 will rotate with the point M on the fixed plane. Therefore, rotational motion of lineN1N2 aroundO2 can be regarded as the synthesis of two movements. One is translational motion of line N1N2 along x2 and y2, and the other is rotational movement round point M. As shown in Fig. 3, taking any point N in the straight line N1N2, and supposing the translational velocity of point M is vM, the translational velocity of point N is vN in instantaneous. The plane MN1N2 rotates around the point M, and its angular velocity is \u03c9T. The generating plane rotating around the gear axis O2, \u03c92 means the theoretical angular velocity, when line segments O2M, O2N and MN are represented by vectors rM, rN, and rNM, respectively. The velocity can be expressed as: vN \u00bc vM \u00fe \u03c9T rNM \u00f08\u00de where vN = \u03c92 \u00d7 rN and vM = \u03c9M \u00d7 rM, while \u03c92 \u00d7 rN =\u03c9M \u00d7 rM +\u03c9T \u00d7 rNM" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002433_j.ijsolstr.2020.03.020-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002433_j.ijsolstr.2020.03.020-Figure3-1.png", "caption": "Fig. 3. The overlapping columns system corresponding to the structural model depicted in Fig. 1 -a).", "texts": [ " The first four coefficients p i of expression ( 8 ) thus comuted are: p 0 = K A ( 1 \u2212 2 L u path, 2 )2 + K B L ( 2 L u path, 2 \u2212 1 ) = P crit p 1 = 3 L u path, 3 [ 2 K A ( 1 \u2212 2 L u path, 2 )2 \u2212 K B ] ( 1 \u2212 2 L u path, 2 )2 p 2 = 1 6 L ( 2 L u path, 2 \u2212 1 )3 { K B \u2212 K A ( 2 L u path, 2 \u2212 1 )3 [ 8 L 3 ( 8 u 3 path, 2 \u2212 9 u path, 4 ) + 1 ] + 2 K B L [ 27 L 3 u 2 path, 3 + 12 L 2 u path, 4 ( 1 \u2212 2 L u path, 2 ) + 8 Lu 2 path, 2 \u2212 5 u path, 2 ]} p 3 = \u2212L 2 ( 1 \u2212 2 L u path, 2 )4 { 4 K A ( 1 \u2212 2 L u path, 2 )4 { u path, 3 [ L u path, 2 ( 32 L u path, 2 + 1 ) + 1 ] \u2212 10 L 2 u path, 5 } + 3 K B u path, 3 { 2 L 2 [ 9 L 2 u 2 path, 3 + 8 L u path, 4 ( 1 \u2212 2 L u path, 2 ) + 2 u 2 path, 2 ] \u2212 1 } + 10 K B L 2 u path, 5 ( 1 \u2212 2 L u path, 2 )2 } (9, a-d) In expressions ( 9 ) it is worth to highlight that: i) given the value of p 0 , equal to P crit , the post-buckling path emerges obviously from the critical state given by ( \u03b8B , P ) crit = (0, P crit ); ii) by analysing p 1 it is concluded that the bifur- cation is symmetrical when u path ,3 = 0 or when 2 K A (1 \u2212 2 L u path ,2 ) 2 \u2212 K B = 0; V V \u03b8\u23a7\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 t 1 a H H l P a i e t i t L t c a t o i i u > i s 2 t iii) for the symmetrical bifurcation case, positive values of p 2 provoke a stable post-buckling behaviour, and unstable oth- erwise; iv) the values of the stiffness coefficients K A and K B are not commutative, so that swapping springs A and B leads to distinct stability behaviours. 2.3.1. Definition of the corresponding structural system, its energy formulation and critical load analysis It was noticed just above that the critical behaviour of the structural system depicted in Fig. 1 is given by expression ( 4 ) and depends on the curvature of the support path at point A, which corresponds to the inverse of the radius of a circle tangent to the path at this point, given by expression ( 5 ). Based on this observation, the structural system depicted in Fig. 3 is now proposed. It consists of two rigid rods [AC] and [AB] connected by rotational springs at A and B, and for a tensile force applied at section B, rod [AC] becomes compressed, which causes instability. The analysis of the frame in its non-trivial configuration computes the longitudinal displacements of points A and B as functions of the rotations \u03b81 and \u03b82 of bars [AB] and [AC]: { u A = L 2 ( 1 \u2212 cos \u03b82 ) u B = u A \u2212 L 1 ( 1 \u2212 cos \u03b81 ) = L 2 ( 1 \u2212 cos \u03b82 ) \u2212 L 1 ( 1 \u2212 cos \u03b81 ) (10) The total potential energy of the structural system is as follows: = 1 2 K A ( \u03b82 \u2212 \u03b81 ) 2 + 1 2 K B \u03b8 2 1 \u2212 P \u00b7 u B (11) which is submitted to the following restraint: w A, [ AB ] = w A, [ AC ] \u21d2 G ( \u03b81 , \u03b82 ) = L 1 sin \u03b81 \u2212 L 2 sin \u03b82 = 0 (12) to assure continuity between both bars in the frame\u2019s deformed shape", ", 2012 ) yields the following TPE\u2019s essian matrix along the fundamental path: F P = [ K A + K B + P L 1 \u2212K A L 1 \u2212K A K A \u2212 P L 2 \u2212L 2 L 1 \u2212L 2 0 ] (15) Rendering null the determinant of H FP yields the frame\u2019s critical oad: crit = K A ( L 1 \u2212 L 2 ) 2 + K B L 2 2 L 1 ( L 1 \u2212 L 2 ) L 2 = K A ( L 1 \u2212L 2 L 2 )2 + K B L 1 ( L 1 \u2212L 2 L 2 ) (16) nd its strong similarity with expression ( 4 ) becomes fully explicit: f, in expression ( 4 ), L u \u2032 \u2032 path (0) \u2212 1 is replaced by L 1 \u2212L 2 L 2 , we obtain xpression ( 16 ). Therefore, for buckling purposes the structural sysem depicted in Fig. 3 is perfectly equivalent to the system shown n Fig. 1 -a) as long as the following geometric condition holds beween both frames: 1 u \u2032\u2032 path ( 0 ) \u2212 1 = L 1 \u2212 L 2 L 2 (17) Consequently, from a structural stability point of view the strucural system of Fig. 1 -a) has a corresponding alternative frame for ritical behaviour, illustrated in Fig 3 , for which tensile buckling rises from the compression of a bar overlapping the one to which he tensile load is applied, in line with the tensile buckling model f Ziegler ( Ziegler, 1977 ). For critical behaviour this correspondence s independent of the springs stiffness coefficients K A and K B and, n line with the statements presented in section 2.2.1 related to \u2032 \u2032 path (0), P crit given by expression ( 16 ) is of tensile type when L 1 L 2 and is of compressive type when L 1 < L 2 , and lim L 2 \u2192 L 1 P crit = \u221e f K B = 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001851_s12206-016-0614-7-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001851_s12206-016-0614-7-Figure17-1.png", "caption": "Fig. 17. Schematic diagram of mechanical properties of leaf spring.", "texts": [ " The dynamic spring rate proposed in this paper reflects the average stiffness of the leaf spring in a cycle of action. It considers that the mechanical properties of the leaf spring are nonlinear and hysteretic, that is, dynamic spring rate is always changing with different load. The calculation of dynamic spring rate has no clear definition; thus, this study proposes a formula to calculate the dynamic spring rate. This paper started from a schematic diagram of mechanical properties of leaf spring, as shown in Fig. 17, to calculate dynamic spring rate. A load-deflection curve can be approximately divided into the segments of F-A-D, D-E, E-C-B, and B-F. The segments of F-A-D and E-C-B represent the normal loading stage and the normal unloading stage, respectively. The segments of D-E and B-F represent the transition from loading to unloading and the transition from unloading to loading. The slope of the curve suddenly increases in both stages when the curve stays in the transitional stage of loading to unloading and unloading to loading", " Dynamic spring rate is a variation, but its changes may be regarded as the process that the value of dynamic spring rate changes from K1 to K2 and then from K2 to K1. Apparently, it is difficult to obtain the specific value of K1 and K2. The dynamic spring rate formula defined herein reflects K1 and K2. Therefore, this paper presents a formula for dynamic spring rate, which is denoted with K, as follows: x FK = V V (11) where DF is the variation of load, and Dx is the variation of deflection. Fig. 17 shows the mathematical implication of K. Eq. (11) indicates that the specific value of dynamic spring rate from simulations and tests can be calculated. The relative error of dynamic spring rate is then obtained. Table 2 shows the relative error of dynamic spring rate. The relative error of dynamic spring rate from the modified model can be controlled within 5%, that is, the precision of the modified model is about 95%. The relative error of the uncorrected model is excessively large to reliably represent the mechanical properties of the taper leaf spring of tandem suspension" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure9.16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure9.16-1.png", "caption": "Figure 9.16 Locating workpiece on faceplate", "texts": [ "11 (a) Outside and (b) inside jaws The four-jaw independent chuck, Fig.\u00a09.12, is used to hold square, rectangular and irregular-shaped work which cannot be held in the three-jaw selfcentring type. It is available in sizes from 150 mm to 1060 mm. As the name implies, each jaw is operated independently by means of a screw \u2013 the jaws do not move simultaneously. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 9 135 9 Turning hold the workpiece against the faceplate while clamping is carried out, Fig.\u00a09.16. Accidents occur when chuck keys are left in the\u00a0chuck and the machine is inadvertently switched on. No matter for how short a period, never leave the chuck key in the chuck. Safety chuck keys, Fig.\u00a09.15, are now available which are springloaded and, if left in position, pop out and fall from the chuck. Workpieces which already contain a hole which is to be enlarged, e.g. cored holes in a casting, can be marked out to produce a box in the correct position, the sides of which are the same length as the diameter of the required hole", " These consist of a hardened and ground steel bush of known diameter with the ends ground square and a flanged screw. The required hole positions are marked out and a hole is drilled and tapped to suit the screw. The accuracy of the drilled and tapped hole is not important, as there is plenty of clearance between the screw and the bore for the button to be moved about. The button is then held on the work by the screw and is accurately positioned by measuring across the outside of adjacent buttons using a micrometer. The buttons Figure 9.15 Safety chuck key The faceplate, Fig.\u00a09.16, is used for workpieces which cannot be easily held by any of the other methods. When fixed to the machine, the face is square to the machine-spindle centre line. A number of slots are provided in the face for clamping purposes. Workpieces can be clamped to the faceplate surface but, where there is a risk of machining the faceplate, the workpiece must be raised from the surface on parallels before clamping. Positioning of the workpiece depends upon its shape and the accuracy required. Flat plates which require a number of holes are easily positioned by marking out the hole positions and using a centre drill in a drilling machine to centre each position" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003537_10775463211013922-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003537_10775463211013922-Figure1-1.png", "caption": "Figure 1. Schematic of recirculating linear roller bearing guideway.", "texts": [ " In Section 3, the friction forces are described using the proposed modified LuGre model, and justifications for LuGre model modification are presented in detail. Experimental test setup and observed behavior of the roller guideway system are provided in Section 4. The experimental and simulation results are compared in Section 5, demonstrating the proposed model\u2019s superiority over the existing ones in predicting the friction forces. Finally, some conclusions are made in Section 6. The considered recirculating linear roller bearing guideway system is shown in Figure 1; it accommodates predictable and repeatable positioning accuracy. In this study, its behavior is investigated when driven by a zero-mean oscillating velocity. To determine friction (horizontal) and normal dynamic forces, EHL\u2019s solution and asperity line contact of roller-to-raceway in mixed lubrication regime is considered. The hydrodynamic pressure distribution and the film thickness in EHL are obtained using the Reynolds equation, considering the surface deformations. In the mixed lubrication regime, the load in rough EHL contact is shared between the lubricating film and asperities" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002659_01691864.2020.1803128-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002659_01691864.2020.1803128-Figure1-1.png", "caption": "Figure 1. Centroidal motion model of humanoid robot.", "texts": [ " These subproblems cannot be discussed in a completely decoupledmanner, however, because realizable foot placement, centroidal motion, and whole-body motion are all restricted by the kinematics and dynamics of the robot. How to efficiently handle this interdependency of these subproblems is still an open problem. Some studies propose to solve centroidal andwhole-body planning problems simultaneously ( e.g., see [16]). In the following sections, some notable methods are reviewed with primary focus on centroidal trajectory generation. The derivation of the centroidal dynamics of humanoid robots is reviewed in the following.More detailed derivation can be found in [17]. See Figure 1 for illustration. Consider a Cartesian coordinate system whose z-axis is the vertical axis. The ground plane is defined as z = 0. Generally, the net contact force that acts on a robot can be represented by a wrench [f T, \u03c4T]T, where f , \u03c4 are the translational and rotational components of the wrench, respectively. The centroidal equation of motion is expressed as mp\u0308G = f \u2212 mg (1a) L\u0307G = \u03c4 \u2212 pG \u00d7 f (1b) where pG denotes the center-of-mass (CoM), LG denotes the total angular momentum around the CoM, m is the total mass of the robot and g = [0, 0, gz]T is the acceleration of gravity" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000834_978-3-319-16190-7_4-Figure4.3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000834_978-3-319-16190-7_4-Figure4.3-1.png", "caption": "Fig. 4.3 Different stages of rotation for Bryant angles. a Initial global system coordinate; b First rotation; c Second rotation; d Third rotation", "texts": [ " The first rotation may be carried out counterclockwise about the x-axis through an angle \u03d51. The resulting coordinate system is labeled \u03be\u2019\u2019\u03b7\u2019\u2019\u03b6\u2019\u2019. Then, a second rotation, through an angle \u03d52 counterclockwise about the \u03b7\u2019\u2019-axis, produces the intermediate coordinate system \u03be\u2019\u03b7\u2019\u03b6\u2019. Finally, the third rotation, counterclockwise bout \u03b6\u2019-axis through an angle \u03d53, results in the \u03be\u03b7\u03b6 coordinate system. The angles \u03d51, \u03d52 and \u03d53, which are the Bryant angles, completely specify the orientation of the \u03be\u03b7\u03b6 frame relative to the xyz coordinate system. Figure 4.3 shows the various steps of the sequence of rotations associated with Bryant angles. Similarly to the Euler angles, when using Bryant angles, the elements of the complete rotational transformation matrix A can be obtained as the triple product of the matrices that define the elemental planar rotations, i.e., the matrices (Nikravesh 1988) D \u00bc 1 0 0 0 c/1 s/1 0 s/1 c/1 2 4 3 5; C \u00bc c/2 0 s/2 0 1 0 s/2 0 c/2 2 4 3 5; B \u00bc c/3 s/3 0 s/3 c/3 0 0 0 1 2 4 3 5 \u00f04:4\u00de Hence, the complete transformation matrix, A = DCB, is given as A \u00bc c/2c/3 c/2s/3 s/2 c/1s/3 \u00fe s/1s/2c/3 c/1c/3 s/1s/2s/3 s/1c/2 s/1s/3 c/1s/2c/3 s/1c/3 \u00fe c/1s/2s/3 c/1c/2 2 4 3 5 \u00f04:5\u00de In a similar manner as in the zxz convention, the transformation matrix associated with the Bryant angles is highly nonlinear in terms of the three angles" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001720_ecce.2014.6953594-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001720_ecce.2014.6953594-Figure5-1.png", "caption": "Figure 5. Dimensions and geometry used in the simulations (enclosed end winding region)", "texts": [ " NUMERICAL SIMULATIONS OF END WINDING HEAT TRANSFER CHARACTERISTICS Some preliminary proof of concept three dimensional simulations using COMSOL Multiphysics simulation software (Version 4.3b) have been carried out to compare the proposed cooling approach with the more common air cooling approach for cooling end windings of electric machines. These simulations are based on 2 [kW] of heat dissipation per end of an electric machine (96% efficient 100 kW machine (data provided by ABB Inc.)) with prescribed enclosure wall temperatures of 80 \u00b0C (dimensions are shown in Fig. 5). The thermal boundary conditions for ferrofluid cooling (FF filled enclosure) are shown in Fig. 6. It should be noted that in these simulations only heat dissipation and cooling of the end windings are considered. Note that axial symmetry allows the study of a coil segment. The magnetic boundary conditions for modeling the induced magnetic field in the TSFF around the end winding are presented in Fig. 7. For modeling the thermomagnetic convection of temperature sensitive ferrofluid fluids, the fluid is assumed as a single phase Newtonian fluid" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002110_iros.2016.7759680-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002110_iros.2016.7759680-Figure4-1.png", "caption": "Fig. 4. Defined frames for pose integration.", "texts": [ " The pipeline has three stages. The 3D map is firstly stored in the model view stage. Then the map is integrated with the movements of the camera and the VR device in the projection view stage. Finally the configuration of the VR device for display is applied in the perceptive stage. Before rendering the map to the VR device, the pose of the RGB-D camera in the real world needs to be integrated with the viewpoint in the virtual environment for display purpose. Several coordinates are defined as shown in Fig. 4. The world frame is defined as the initial position of the mapping process in the real world space. The relative transformation matrix GWC between the world frame and the camera frame is obtained from the 3D mapping method mentioned in the previous section. The transformation matrix GWV represents the VR device\u2019s relative motion respect to the world frame, which is calculated using the data (roll, pitch and yaw) from the IMU in the VR device. GVE is the projection matrix from the 3D world to a screen or display" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002953_0954406220974052-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002953_0954406220974052-Figure4-1.png", "caption": "Figure 4. The schematic diagram and force analysis of HFFD6. (a) force analysis of a single chain of the parallel structure. (b) force analysis of the moving platform.", "texts": [ " However, the dynamic formula of a parallel robot is complex, and the derivative of the angle and angular velocities need to be calculated also, causing more calculation time to be consumed. For a force feedback device, a considerable high-frequency force refresh rate is crucial to display the realistic haptic experience, usually reaching 1KHZ. Considering the characteristic of the manipulating process, that is, the speed is slow and stable, the inertial force term M\u00f0h\u00de and the centrifugal force term V\u00f0h; _h\u00de can be negligible compared with the gravity term G\u00f0h\u00de. Hence, the active gravity compensation can be simplified into the static balance of HFFD-6 in its workspace. Figure 4 (a) shows a single chain of the 3-DOF parallel mechanism. When the torque is analyzed, the serial mechanism can be approximately simplified into a mass point GF, which acts at the center of the moving platform. As shown in Figure 4(a), the passive linkage Lb is only affected by the forces exerted by the moving platform and the active linkage La. Therefore, the passive linkage is balanced under the action of two forces, and the internal force of passive linkage is recorded as f i. In the coordinate frame {O}, the unit vector nBiCi can be expressed as follows nBiCi \u00bc r Lasinhi R\u00f0 \u00decosgi \u00fe x Lb r Lasinhi R\u00f0 \u00desingi \u00fe y Lb z Lacoshi Lb 2 666666664 3 777777775 gi \u00bc 4i 1\u00f0 \u00de 6 p; i \u00bc 1; 2; 3 (9) where the coordinates of the center of the moving platform P\u00f0x; y; z\u00de can be calculated according to the forward kinematics shown in equation (3). In the case of active gravity compensation, as shown in Figure 4(b), the moving platform maintains balance under the combined force of the passive linkage f i, its gravity GM, and one-third of gravity of passive linkage GP, which is recorded as GP1. The forces on the moving platform can be expressed as the matrix f \u00bc f1 f2 f3 T , and it can be obtained by analyzing its force in x; y; z directions. f1 f2 f3 2 664 3 775 \u00bc H 1 0 GM \u00fe 3 GP1 0 2 664 3 775 (10) where H \u00bc A 1 A 2 A 3 B 1 B 2 B 3 C 1 C 2 C 3 2 664 3 775 A i \u00bc r Lasinhi R\u00f0 \u00decosgi \u00fe x Lb ; B i \u00bc r Lasinhi R\u00f0 \u00desingi \u00fe y Lb ; C i \u00bc z Lacoshi Lb As shown in Figure 4(a), the force that passive linkage Lb exerts on the active linkage La is recorded as f i, and f iI is the decomposed force of f i in the plane OAiBi, f iP donates the decomposed force of f i perpendicular to the plane OAiBi. Note that the force perpendicular to the plane OAiBi doesn\u2019t work. From the relationship f i f iP \u00bc f iI\u00f0i \u00bc 1; 2; 3\u00de (11) Combine with equation (9), f i can be expressed as follows where f iP \u00bc f i nci, and nci donates the unit normal vector of the plane OAiBi. To simplify the calculation, since the gravity of the passive linkage is distributed equivalently, so two-third of its gravity is assigned to the active linkage and is recorded as GP2 GP2 \u00bc 0 2 3GP 0 h iT (13) The decomposed force of GP2 in the plane OAiBi is recorded as GI P2, and the one perpendicular to the plane OAiBi is recorded as GP P2, where G P P2 \u00bc GP2 nci" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003053_iros45743.2020.9341428-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003053_iros45743.2020.9341428-Figure1-1.png", "caption": "Fig. 1: Planning challenges inherent to the LEGO domain in contrast to other common blocks world examples.", "texts": [ " Bricks can be stacked within this grid using right-angled rotations. As putting two LEGO bricks together can be seen as a joining process [1], the building of LEGO structures is a perfect example for assembly. Due to its combinatorial complexity, the LEGO domain is a suitable representative for automatically planning robotic assembly sequences and for evaluating different approaches and algorithms. In contrast to common blocks world examples for planning [2], LEGO offers additional challenges. Fig. 1a shows a situation in which a LEGO brick can no longer be placed in its target position due to the knobs. This exclusion problem requires to dismantle the structure partially in order to place the missing brick. A related problem occurs when a LEGO brick cannot be placed due to a collision between neighboured bricks and the gripper holding this brick (cf. Fig. 1b). In both situations, a planner can still explore a huge set of further states without detecting this deadlock situation. This requires appropriate strategies for backtracking at an early stage in order to efficiently find a plan. If LEGO bricks are assembled in order to form a staircase, the structure will collapse from a certain height due to its own weight. The staircase depicted in Fig. 1c is only stable if its upper part is supported, e. g., by the other structure shown. If you imagine a manual assembly of such a staircase, every placement of a brick would have to be supported by a second hand until the overall construction reaches a stable condition. Hence, the statics of LEGO structures must be *This work is partly funded by the German Research Foundation (DFG) under the TeamBotS grant. The authors are with the Institute for Software & Systems Engineering, University of Augsburg, Universita\u0308tsstr. 6a, Augsburg, Germany. ludwig.naegele@informatik.uni-augsburg.de regarded during planning. Another challenge, the overhang problem, is shown in Fig. 1d. Here, a straightforward bottomup construction would not lead to a valid plan, since the bricks of the overhanging substructure can only be attached towards the end, although they belong to lower levels. That is why a simple bottom-up strategy \u2013 level by level \u2013 is not feasible in general. Instead, a planning approach is required that is independent of the spatial position of LEGO bricks. In this paper, we propose an AI-based planning approach and its formal description for coordinated multi-robot assembly", " The house has an area of 128mm\u00d7 128mm and a height of approximately 116mm. It consists of a lower base part with one door and three window sides and a pyramid like roof top. The structure is built on a ground plate with 26 bricks on 6 levels: 4 red and 9 yellow bricks with a grid of 2\u00d7 2 knobs as well as 10 red and 3 yellow bricks with 4\u00d7 2 knobs are used. The house can be constructed using one or more robots and contains multiple deadlock situations on the corner bricks of the roof top due to the collision problem depicted in Fig. 1b. Such a deadlock situation is not detectable at the time it originally occurs, because there are still many permissible subsequent steps. This leads to high planning times with exponential complexity subject to the number of bricks and skills. Even this rather simple example might not be solvable in adequate time when having to examine all these possibilities. The second LEGO case study is the DUPLO bridge which is depicted in Fig. 2c. The arch bridge spans approximately 290mm and is built on a ground plate with 64 LEGO bricks on 14 levels: 32 yellow bricks with a grid of 2\u00d7 2 knobs, 22 yellow bricks with 4\u00d7 2 knobs, one yellow brick with 10\u00d7 2 knobs as well as 6 yellow bricks with 2\u00d7 1 knobs and double height are required for the arch. For the road on top, 3 grey bricks with a grid of 8\u00d7 2 knobs are necessary. This case study was chosen because it can only be assembled with at least three robots and it shows all the problems of the LEGO domain introduced in Fig. 1. If the arch bridge is built from one side, the remaining stones on the other side of the arch can be added from the top brick to the bottom and, thus, can be viewed as an overhang. However, this overhang situation would also lead to an exclusion, since at least the last brick between the almost finished arch and the base plate can no longer be inserted. Moreover, the exclusion problem can occur at both abutments of the arch. Below the imposts, there are a large number of possible sequences to place the LEGO bricks in order to form both pillars, which lead to collisions between gripper and bricks", " Analysis Modules identify and extract LEGOspecific types of attributes from a given 3D model. The so-called LegoPlacementAttribute describes that a brick is being stacked on top of another with a given geometric offset. LegoSupportAttribues are used during planning and indicate that a robot is giving additional support to a brick in order to stabilize the construction. Validation Modules reject invalid planning states from further planning: While one LEGOspecific module aims at discovering exclusions as motivated in Fig. 1, another module also considers robots and performs statics analysis in order to determine the stability of a construction. Fragile constructions, which are not adequately supported by robots and, thus, would collapse due to their own weight, are reliably rejected. The planning of micro steps uses Skill Modules to retrieve possible automation tasks that can be performed with robots in a current planning situation. For each robot, a PickAndPlaceSkill is provided that creates pick-and-place tasks in a given situation for each brick that is reachable by the robot" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003001_j.engfailanal.2020.105195-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003001_j.engfailanal.2020.105195-Figure1-1.png", "caption": "Fig. 1. Bicycle frame geometry.", "texts": [ " The development of the procedures includes identification of fatigue properties, cracking resistance in limited volume materials [20,21], verification of local properties of thin-walled structures [22,23] and welded joint areas [24]. The use of miniature specimens also reduces the costs of fatigue tests [25]. The analyses were carried out for a bicycle frame used in gravity mountain biking. The bicycle frame is made of thin-walled tubes made of high-resistance steel alloyed with chromium and molybdenum. Fig. 1 shows the geometry of the analysed bicycle frame. After three years of use, fatigue cracks were detected at the connection between the top tube and the seat stays. The crack was initiated on both sides of the bicycle frame. Fig. 2 shows the failure location. The crack is located near the weld bead, and most of it is in the heat-affected zone. Fig. 3 shows the actual bicycle frame node with a visible fatigue crack. The crack is 10 mm long and is visible on the inner and outer sides of the tube. The crack is not directly on the edge of the welded joint, and it is offset from the edge by 1", " Tomaszewski Engineering Failure Analysis 122 (2021) 105195 A fatigue strength approach using nominal stresses with maximum dynamic and quasi-static loads is a method commonly used to determine the fatigue properties of a structure. The stresses are calculated using a finite element model including surface elements. The highest stress points are selected for comparison of different fatigue evaluations. A finite element analysis in ANSYS Workbench 2020 R1 software was carried out to determine the stress distribution within the analysed object. The geometry of the numerical model is consistent with Fig. 1 and was prepared in a CAD environment based on the data provided by the manufacturer and the measurements of the actual test object. The conditions of the numerical analysis conform to ISO 4210\u20136 [5], developed by a special technical committee TC149 on the standardization in the field of cycles, their components and accessories with particular reference to terminology, testing methods and requirements for performance and safety. The standard includes safety requirements and test methods for bicycle frames for selected loading cases depending on the bicycle type (city and trekking, young adult, mountain, racing)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002286_j.cagd.2020.101826-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002286_j.cagd.2020.101826-Figure1-1.png", "caption": "Fig. 1. Types of impossible figures: a) Depth interposition, b) Depth contradiction, c) Disappearing normals, d) Disappearing space.", "texts": [ " Impossible figures (Ernst, 1986), also called impossible objects (Ernst, 1996; Unruh, 2001) or inconsistent images (Mortensen, 2010), are an optical illusion designed to play with your visual perception. Popularized by the celebrated Dutch artist M.C Escher, these 2D drawings lead to a contradiction when the viewer tries to interpret them as the projection of a 3D object. Adapting previous works (Ernst, 1986; Kulpa, 1987), Mortensen (2010) or Wu et al. (2010) categorize these figures into four classes, exemplified with classical models in Fig. 1: a) Depth interposition: The apparent depth ordering leads to structural inconsistency. Typical examples are the Necker cube, or a four-bar arrangement. b) Depth contradiction: The propagation of local depth information leads to a structural depth contradiction. Examples are the endless Penrose stairs, or the impossible four-bar. \u2729 Editor: Rida Farouki. * Corresponding author. E-mail address: Javier.SanchezReyes@uclm.es (J. S\u00e1nchez-Reyes). https://doi.org/10.1016/j.cagd.2020.101826 0167-8396/\u00a9 2020 Published by Elsevier B", " Remarkably, this reaction of \u201csurprise\u201d (or \u201cinterest\u201d) when viewing impossible figures has been detected in infants from a very early age (Krause et al., 2019). Nevertheless, dropping the assumption of a generic view opens up the possibility of creating 3D impossible objects. The trick boils down to designing a model such that, if seen from a very specific viewpoint, this accident of the viewpoint results in a violation of a rule and hence in a non-generic view. For instance, consider the impossible four-bar of Fig. 1b, thoroughly analyzed in the literature (Tsuruno and Tomimatsu, 2017) and often used as a representative impossible figure (Chiba et al., 2018), since its inception goes back to 1568 (in a painting by Brueghel). If seen from the accidental vantage point O (Fig. 2a), the four-bar can be materialized with two different models: (1) S (discontinuous, polyhedral): Yet it meets assumption (2), it violates (1). The closed projection fools the viewer into thinking that S is also closed, leading to depth contradiction", " Perspective projection is trickier, as it does not preserve parallelism, so we interpret several edges as parallel when they meet at a common vanishing point V outside the drawing (Fig. 5b). Also by grouping lines converging to vanishing points (Company et al., 2014), automated reconstruction methods interpret sketches of polyhedra in perspective projection. Sugihara (2018) has resorted to the violation of this rule in his award-winning impossible motion objects. We will also use this technique later to materialize the Renault logo (Fig. 1c). Strictly speaking, the expression look parallel in rule (3), and similarly for straight in rule (2), should be interpreted in the setting of this section, i.e., from a perceptual viewpoint. Our ability to perceive straightness and parallelism is an elaborate issue, whose accuracy we cannot take for granted, as Rogers and Naumenko (2016) warn. The observer\u2019s inability to perceive depth is the key feature impossible objects exploit. Indeed, in the example of Fig. 2, take a point p on the plane where the impossible figure is drawn, and then move p along a radial direction Op through the viewpoint, in what Elber (2011) calls Line of Sight Deformation, to new locations p or p\u0302", " The models S\u0302 contain two slender through holes, materializing rays fired from the viewpoint, to help the viewer orient and position S\u0302 correctly. The viewer aligns S\u0302 so that one hole is seen through, and then adjusts the distance until the second hole is also seen through. For better effect, the model must be viewed with only one eye, to avoid any depth perception resulting from binocular vision. Defining non-planar rectangular patches with straight boundaries is a simple task in the NURBS model, as explained in Section 3.3. We use this technique to materialize the Renault logo (Fig. 1c), designed by the Hungarian-French graphic artist Victor Vasarely. First, we construct a polyhedral version S , as in the previous case, in three steps: 1- Create an extruded rhombus, made up of four planar facets. 2- Find the point p , on the projection plane , where the projections of the edges e (blue) and g (red) intersect, and compute the corresponding point p on e. 3- Using the trimming line pq, take the rectangular facet displayed in blue and trim away the upper triangular area, which furnishes a quadrilateral F ", "2), as the new edge e\u0302 is no longer parallel to the original edges of the rhomb. The texture, with stripes parallel to e, reinforces this parallelism effect, tricking the viewer into interpreting F\u0302 as a (partially occluded) parallelogram, which leads to an inconsistent normal. Mathematically, the orientable surface S has transformed into its non-orientable anamorphic counterpart S\u0302 , a G0 homeomorphic analog of a Moebius strip. This topology supports the classification as an impossible object with disappearing normals (Fig. 1c) In Fig. 12, much like in Fig. 3, we have shifted the position of the viewpoint, from the accidental position O, and then displayed the resulting views for both S and S\u0302 . Once again, the deception still works only for the twisted version S\u0302 , mostly thanks to the texture, whereas in S it highlights the edge mismatch, thereby revealing the trick. We have brought together concepts from different areas (theory of perception, anamorphic art, 3D reconstruction of line drawings, computer graphics, and CAGD) usually overlooked in the literature on impossible figures" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000754_1.4031440-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000754_1.4031440-Figure2-1.png", "caption": "Fig. 2 Prototype oil-free TC supported on gas bearings (CAD model and photograph of test hardware)", "texts": [ " Furthermore, extensive experiments are conducted in a gas stand test facility to compare power loss and system efficiency between the conventional floating ring bearing supported TC and the gas bearing supported TC which incorporates the same compressor (CW) and turbine wheel (TW). Presently, two TCs (namely, Core 1 and Core 2) are built; Core 1 and Core 2 are supported on foil journal bearings and FPTP journal bearings, respectively. Except journal bearings, other components in Cores 1 and 2 are identical. That is, both TCs use the same bearing (inner and outer) housings, foil thrust bearings, rotor, TW ( 50 mm in diameter), and CW ( 55 mm in diameter) with identical geometries and mechanical properties. Figure 2 shows one of the test oil-free TCs with steel main body (i.e., bearing outer housing) integrating an aluminum water jacket. The water jacket (i.e., bearing inner housing) includes a cooling water passage, extending circumferentially about the bearing sleeves, for circulating cooling water through the BH. Note that, currently, no forced air flow streams are supplied into the gas bearings.1 Figure 3 illustrates a schematic view (not to scale) for the dimensions of the test rotor, the bearings, and BH" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure9-1.png", "caption": "Fig. 9. Position of loading devices.", "texts": [ " 6-DOF loadings are possible using loading devices located in the axial and radial directions, and revolutions in both directions are also possible using the drive system. Radial (horizontal) direction loading is applied to the outer ring of the test bearing, whereas axial (vertical) direction loading is applied to the plate that is connected to the inner ring of the auxiliary bearing. The rig can also be used to test internal teeth type bearings by changing the position of the driving system and insert plate. From the bearing\u2019s center point, four vertical and two horizontal positions are selected as loading points to apply 6-DOF loadings (Fig. 9). By using the force balancing condition, it is possible to determine the loading device\u2019s loads\u2014A1, A2, A3, A4, A5, and A6\u2014to reproduce each external loading\u2014Fx, Fy, Fz, Mx, and My\u2014as shown in Eq. (1). These loads are specified to compensate the weight of the plate on the bearing, W. 11 ( ) 2 1 12 ( ) 2 1 3 4 2 MxA Fz W L MxA Fz W L A Fy MyA L = + + = + - = = 5 2 6 . MyA L A Fx = - = (1) The final test rig was developed using the loading conditions and bearing size information obtained from a manufacturer of 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.18-1.png", "caption": "Fig. 3.18 A nested or cascaded rotation can, for example, be the orientation of a camera [cxcycz] on a moving base [bxbybz] fixating a target, here the top of a mountain", "texts": [ " For example, the moving eye is placed inside a moving head, or a gimballed camera is mounted on a moving vehicle or missile. How can the formulas given above be used to derive the composite rotation of the two3-D rotations combined? Furthermore, given the location of a point in space-fixed coordinates and eye coordinates, would it be possible to calculate the required rotation to ensure that the eye/camera looks at the point? To describe the orientation of a camera-in-space (described by Rspace camera), as shown in Fig. 3.18, one has to combine the orientation of the tilted base, e.g., a Google maps car (described by Rspace base ) and the orientation of the camera with respect to this base (Rbase camera). To implement this mathematically, one has to use rotations of the coordinate system. This determines the sequence of the two rotations, and the rotation matrix describing the orientation of the camera-in-space is\u2013according to the 3.6 Applications 53 discussion following Eqs. (3.20) and (3.22)-given by Rspace camera = Rspace base \u00b7 Rbase camera " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002672_ccdc49329.2020.9164017-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002672_ccdc49329.2020.9164017-Figure1-1.png", "caption": "Fig 1. Three-dimensional model of dulcimer music robot", "texts": [ " Therefore, it can effectively solve the trajectory optimization problem of robots under kinematic constraints. In this paper, for an anthropomorphic music robot of playing dulcimer, an improved PSO algorithm is proposed and used to optimize the total time of the trajectory planning under complete kinematic constraints. This will greatly help to improve the performance of the dulcimer music robot. The dulcimer music robot in this paper consists of a mechanical arm, a wrist striking device, a support base and a dulcimer. The three-dimensional model is shown in Fig. 1. According to the manipulator configuration of dulcimer This work is supported by the Planning Fund of the Department of Social Sciences of the Ministry of Education under Grant 16YJAZH080. music robot, the kinematic model of the robot arm is established by D-H parameter method [10]. Because the structures of the left and right arms are same, the left arm is modeled in this paper, the D-H link coordinate system is shown in Fig. 2, and the link parameter of left arm is shown in Table 1. 4769978-1-7281-5855-6/20/$31" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000166_01691864.2019.1608299-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000166_01691864.2019.1608299-Figure11-1.png", "caption": "Figure 11. Schematic view of the experimental setup for the grasping and releasing test (refer to the video clip).", "texts": [ " The efficacy of the proposed surface composed by the slit texture to obtain a high friction and the lubricating system for reducing friction was examined in the previous section. In this section, the developed friction reduction system was evaluated at the grasping and releasing operations. Here, supposing that the tasks are performed in a narrow space, we investigated whether a grasped object can be released by controlling the friction, without changing the grasping configuration or pose. The target objects are shown in Figure 10 while the experimental setup is shown in Figure 11. A gripper in which the surface corresponds to the developed surface was utilized. The procedure is as follows. First, the target object on the stage was grasped by the gripper without injecting the lubricant with the grasping force listed in Table 5. Subsequently, the stage was removed and the weight with the value listed in Table 5 (approximately 30\u201340% that of the target object) was placed on the top of the object to confirm the realization and stability of grasping. After removing the weight, the lubricant was injected without changing the grasping configuration to confirm friction reduction by examining whether the object slips down" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000566_ab5b65-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000566_ab5b65-Figure13-1.png", "caption": "Figure 13. The worn surfaces with interval of 13 teeth on the planet gear.", "texts": [ " For comparison, the result from original RPCA suffers from the residual interferences, which shows the advantage of PRPCA on extracting target fault signatures. Surface wear of the tooth is another common failure contributing to the insufficient fatigue strength. In the manufacturing industry, surface wear will affect machining accuracy, which may cause the scrap of parts. In this case, two teeth with an interval of 13 teeth on the gear are artificially worn to validate the applicability of the proposed method. The gear is shown in figure\u00a013. Figures 14(a) and (b) present the results of TSA and MTSA, from which we can find that MTSA reserves much more information about the gear meshing no matter on the amplitude or the regularity. Then, PRPCA and RPCA are both used to extract the anomaly as a comparison. Theoretically two distinct impulses are expected to appear within each cycle, however, this is not completely achieved by RPCA from figure\u00a014(c) where interferences disturb the recognition of fault impulses of small amplitude. By embedding the local periodicity, PRPCA presents its superior ability on extracting complete fault information, which can be seen in figure\u00a014(d)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003324_j.mechmachtheory.2021.104261-Figure22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003324_j.mechmachtheory.2021.104261-Figure22-1.png", "caption": "Fig. 22. Hexacopter with the deployable 2-dof parallel manipulator.", "texts": [ " By utilizing the light tension components, the mechanism can perform good stability without large weight gain. All the experiment results are consistent with the theoretical conclusions in the previous sections. The 2-dof DTPM is collapsible for easy storage, and can be deployed for a variety of operations. This kind of manipulator is suitable for using on small platforms, e.g. UAV, for the purpose of pick-and-place, grasp and inspection. In this subsection, the prototype is mounted under a Hexacopter(as shown in Fig. 22 ) and taken into the test flight. In the test flight, the 2-dof DTPM is folded when the Hexacopter is taking off ( Fig. 23 (a)). After the Hexacopter flies to a certain height, the manipulator starts to extend ( Fig. 23 (b)). By controlling the two rotary actuators, the gripper at the tip of the manipulator can translate in its local OXY plane to perform tasks ( Fig. 23 (c)). To prepare for landing, the manipulator is retracting. After it is fully folded, the Hexacopter can land safely ( Fig. 23 (d))" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000411_s11665-019-04291-w-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000411_s11665-019-04291-w-Figure2-1.png", "caption": "Fig. 2 Schematic diagram of the experimental setup", "texts": [ " Once freezing has completed, the samples were freeze-dried in the freeze-dryer for at least 48 h to ensure the complete removal of ice crystals. Rough samples were sintered in a vacuum furnace equipped with a digital temperature control system at a pressure of 4.0 9 10 4 Pa. The sintering temperatures were varied from 800 to 1100 C with a 50 degree interval. The heating rate was set at 3 C/min. Once the target temperature reached, it was hold for 6 h, followed by cooling at a rate of 5 C/min to room temperature. The schematic diagram of the experimental setup is shown in Fig. 2. The porosity was calculated using the mass and dimensions of the specimens in the following equation: P \u00bc 1 mTi=VTi qTi 100 where P is the total porosity percentage, qTi is the theoretical density of titanium, mTi is the mass of the specimen, and VTi is the volume of the specimen. The surface pore structures of the Ti scaffolds were characterized by scanning electron microscopy. The internal pore structure was examined using a microcomputed tomography, (lCT) unit, Skyscan 1272. The scaf-Fig. 4 3D image of the Ti scaffold after sintering at 1000 C Journal of Materials Engineering and Performance folds were positioned inside a tube and imaged with 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure1.11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0003349_s00170-021-06813-0-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003349_s00170-021-06813-0-Figure9-1.png", "caption": "Fig. 9 CAD model of active cooling prototype including (1) diaphragm pump, (2) Peltier cooler, (3) LM reservoir, (4) in situ SLS cold plate with embedded copper-plated ceramic tubes, (5) MP9100 Caddock resistor (12\u00d7), (6) nylon coupling, (7) polypropylene flex tube, and (8) stainless steel #4\u201340 mounting hardware (4\u00d7)", "texts": [ " Six in situ AM SLS cold plates were fabricated. During the thermal test, their performances were compared to cold plates fabricated by the conventional assembly technique. The AM cold plates are seen in Fig. 8. Note that cold plate has form factor similar to base plate for 6U CubeSat, enabling mounting two stacks of 1U CubeSat adjacent to each other. In our prototype designated for thermal experiment, heat sources would be mounted to the right side of the block, separated from the left side that is mounted to the Peltier cooler (see Fig. 9). Active cooling prototype is represented by following equations that govern incompressible fluid flow of LM and heat transfer: The continuity equation (conservation of mass) \u2207:V \u00bc 0 \u00f04\u00de The momentum equation (Navier\u2013Stokes equations) \u03c1 \u2202V \u2202t \u00fe V :\u2207V \u00bc \u2212\u2207p\u00fe \u2207:\u03c4\u2212\u03c1g \u00f05\u00de Conservation of energy (the temperature distribution): \u03c1Cp \u2202T \u2202t \u00fe V :\u2207T \u00bc \u2207: k\u2207T\u00f0 \u00de \u00fe Q \u00f06\u00de where V is velocity of LM, \u03c1 is density of LM, T is temperature, p is pressure, t is cycle time, \u03c4 is deviatoric stress, g is gravity vector, \u03b2 is volumetric expansion coefficient, Cp specific heat at constant pressure, k thermal conductivity, Q is volumetric heat source" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001472_s0263574714002458-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001472_s0263574714002458-Figure4-1.png", "caption": "Fig. 4. The cross-section csi of section i and its polar coordinate system.", "texts": [ " , \u03ban, \u03c6n, sn)T . The section i, i = 1, 2, . . . , of a continuum manipulator is bounded by two planes Hi\u22121 and Hi , as shown in Fig. 3. Hi is the plane containing pi with normal zi+1 and shared by section i and section i + 1. For each section i of the manipulator with a nonzero curvature, let plane Pi contain the section i\u2019s circle, i.e., ciri , then the cross-section of section i by Pi , denoted by csi , is a fan-shaped planar region with a width 2wi , bounded by two rays Li,1 and Li,2. As shown in Fig. 4, using a polar coordinate system (\u03c1, \u03b8) with circle center ci as the pole and xi as the polar axis, the region csi can be described https:/www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574714002458 Downloaded from https:/www.cambridge.org/core. Georgetown University Library, on 13 Feb 2017 at 18:46:01, subject to the Cambridge Core terms of use, available at easily by bounds on \u03c1 and \u03b8 as: ri \u2212 wi \u2264 \u03c1 \u2264 ri + wi, (1) \u03b8min i \u2264 \u03b8 \u2264 \u03b8 max, i (2) where, if \u03bai > 0, then \u03b8min = 0 and \u03b8max = si\u03bai ; else, \u03b8min = \u03c0 + si\u03bai, \u03b8max = \u03c0 ", " Since N n (as it requires more than one bounding volume for each arm section), and n1 n, MN + mn1 (M + m)n. Thus, the CD-CoM Algorithm has a far lower order of worst-case time complexity than OPCODE, and we can further show that O(MN + mn1) O[(M + m)n]. https:/www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574714002458 Downloaded from https:/www.cambridge.org/core. Georgetown University Library, on 13 Feb 2017 at 18:46:01, subject to the Cambridge Core terms of use, available at Arm model Algorithm Config. 4 Config. 5 Config. 6 Exact CD-CoM <1 ms 6 ms 10 ms Mesh 1 OPCODE 7 ms 11 ms 16 ms Mesh 2 OPCODE 37 ms 47 ms 53 ms For checking collision of a path of k arm configurations, the worst-case time complexity of CD-CoM is O[k(M + m)n]; whereas, the worst-case time complexity of OPCODE is O[k(MN + mn1) + Nk], with the additional term Nk reflecting the time complexity of refitting the bounding volume hierarchy of the manipulator for each change of configuration. Thus, k(MN + mn1) + Nk k(M + m)n, and moreover, O[k(MN + mn1) + Nk] O[k(M + m)n]", ", min(\u03c11, \u03c12) > ri + wi and \u03b81, \u03b82 both satisfy (or both do not satisfy) (2) then compute the distance between circle center ci and li and obtain point q = (\u03c1q, \u03b8q) on li if the distance is shorter than ri + wi and \u03b8q satisfies (2) return Collision \u2190True else return Collision \u2190 False Case 4: if both \u03b81 and \u03b82 satisfy (2) then return Collision \u2190 True / * the remaining collision cases have intersections between rays Li,1 (or Li,2) and li * / Case 5: if line segment li intersects rays Li,k (k = 1, 2) of csi at pk int (see Fig. 4) then if one vertex of li satisfies (2) and is above the upper bound for \u03c1. then if \u03c1k int \u2264 ri + wi then return Collision \u2190 True (see Fig. 5(e)) end if end if if one vertex of li satisfies (2) and is below the lower bound for \u03c1. then if \u03c1k int \u2265 ri \u2212 wi then return Collision \u2190 True (see Fig. 5(f)) end if end if if neither vertices of li satisfies (2) then if one of \u03c1k int satisfies \u03c1k int \u2265 ri \u2212 wi then return Collision \u2190 True (see Fig. 5(g)) end if end if return Collision \u2190 False. https:/www.cambridge" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002357_robosoft48309.2020.9116040-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002357_robosoft48309.2020.9116040-Figure2-1.png", "caption": "Fig. 2. CAD rendering of the REBO Hopper. The design is an inverted version of the REBO Juggler [30].", "texts": [ " REBO mechanisms exhibit high resilience and tolerate repeated loading and unloading over thousands of cycles with little fatigue and no signs of physical damage, encouraging their use as a lightweight, ideal spring. The parameters for the double-layer REBO spring used in this research are given in Table I, where each REBO spring exhibits a stiffness of Ks = 1035 N m\u22121. By arranging three of these double layer REBO springs in parallel, we get a compliant leg with stiffness of roughly Kes = 3105 N m\u22121. The REBO Hopper robot shown in Fig. 2 uses the REBO bellows for energy storage and release during hopping. The design is a modification of the REBO Juggler in our previous work [30], where we turned the robot upside-down such that instead of pushing a ball into the air, the new robot pushes itself off the ground. The REBO Hopper comprises three major subsystems: (a) the compliant leg, (b) a toe sensor, 5The notation of the pattern has been changed from the previous work [30] to make the parameter more clear. The subscript (i or o) of the cone angle \u03b2, the length parameter a, b and h indicates the inner and the outer layer of the double-layered REBO structure", " As a result, at ground impact, the motors overshoot ptr with a higher time delay. The stance mode of the robot is about 60 ms, and there is enough time for the robot to detect contact using proprioception even with the lowest drop energy tested in this experiment. Using a P-loop controller for proprioceptive measurements (as seen in Eq. 1 of [45]) and similarly written as I\u03b8\u0308 = kp(\u03b8des \u2212 \u03b8) (2) works well in both the drop test and the tendon losing tension. In Fig. 6(b) there is a similar trend to those seen in Fig. 2(a) from [45] and in fact, the time delay seems to improve for the tendon driven system. While the improvement is useful, it is not in the scope of this work to directly compare the magnitudes of energy and time between the two types systems. We conduct the vertical hopping test using the proprioceptive contact detection as the mode guard. The state machine is shown in Fig. 7(b), similar to Fig. 4(b). We have chosen ptr = 0.3 mm as our trigger for contact detection. When the hopper is in flight, the REBO actuator is in the compressed mode where the springs are set to a pre-compressed length ppc" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002254_mawe.201800221-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002254_mawe.201800221-Figure2-1.png", "caption": "Figure 2. Schematic diagram of laser cladding.", "texts": [ " The pure niobium carbide powder is massive and has extremely fine particle size and the particle sizes of niobium carbide are 1 \u03bcm\u20136 \u03bcm, Figure 1b. The chemical compositions of Ni60 A powder are as followings: 0.83 % carbon, 16.52 % chromium, 3.27 % boron (B), 3.64 % tungsten (W), 6.62 % iron (Fe), and Ni: Bal. The laser cladding equipment system used in this experiment is mainly composed of YLS-6000 fiber laser, powder feeding equipment and water cooling system. Under the action of protecting gas and cooling water, laser as heat source are fused synchronously with powder to form coating on the surface of substrate, Figure 2 [24, 25]. According to the control variable method, 11 different process parameters are designed in this experiment. No. 1\u2013 3 controls the same scanning speed and powder feeding rate to increase the laser power. No. 4\u20136 controls the same laser power and powder feeding rate, and changes the scanning speed; No. 4, 9 and 11 control the same laser power and scanning speed, and change the powder feeding rate. The effects of various parameters on the cladding layer were observed, Table 1. The Cr12MoV steel plate was cut into 100 mm\u00d750 mm\u00d710 mm, and the surface rust was removed by an angle grinder, and then laser cladding was performed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001020_s00170-015-7033-2-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001020_s00170-015-7033-2-Figure2-1.png", "caption": "Fig. 2 The double-circular-arc cutting edge", "texts": [ " The double-circular-arc torus milling cutter has high metal removal rate at low depth of cut and high feed rate. The feed rate can be four times or more of the regular torus milling cutter. As the double-circular-arc torus milling cutter has good geometric and cutting performance, to provide theory guidance the fabrication and application of the milling cutter, it is necessary to build the mathematical model of the milling cutter. Fig. 3 The sub-circular-arc cutting edge The double-circular-arc cutting edge is shown in Fig. 2. To show a clear structure of the cutting edge, the subcircular-arc is shown in Fig. 3. In Figs. 2 and 3, O1 and O3 are the centers of main-circular-arc cutting edges; O2 is the center of the sub-circular-arc cutting edge, and the radius is r; and \u03c8 is the rotation angle of cutting edge on jok plane. I n F i g . 2 , |O 1O 3 | = D 1 , |AG | = D 2 , |GH | = a p , R1=|O1B|=|O1C|=|O3D|. Simultaneously solve Eqs. (1) and (2), the radius of the main-circular-arc cutting edge is: R1 \u00bc DG2 2ap \u00bc D2\u2212D1\u00f0 \u00de 2 2 \u00fe a2p 2ap \u00bc D2 \u2212D1\u00f0 \u00de2 \u00fe 4a2p 8ap \u00f015\u00de FromEq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure16-1.png", "caption": "Figure 16. Mesh more refined (Nodes 185901, Elements 113367)", "texts": [ " According to Table 7, one notices that the equivalent maximum constraints of Von Mises increase, according to the number of elements of the grid. The maximum value of the equivalent constraint of Von Mises as well as the total deformation attacks corresponds d to the maximum of the elements of the grid are practically those which one meets in the literature. It is thus judicious to choose a refined grid, because the solution becomes more exact by increasing the number of nodes of the grid. a) Influence smoothness of the mesh For that, one considered a second type of mesh finer and refined in the friction tracks, (Fig.16) appears. The element used in this mesh is a SOLID 187 and the total time of the simulation is equal to 8331.328 (s). This new mesh (type M2) consists of 11 3367 elements TE with 4 nodes, that is to say 18 5901 nodes. It is thus much finer than the mesh M1, Fig. 15 (c) used up to that point. Figure 17 shows the various configurations of displacements of the order \u00ecm of the model according to time, while keeping the symmetrical form compared to the vertical median plane. The total deformation is reached at the end of braking and varies between 0 to 52,829 \u00ecm ", " The initial temperature of the disc and pads is 20\u00b0C, and the surface convection condition is applied to all surfaces of the disc and the convection coefficient (h) of 5 W/m2\u00b0C is applied at the surface of the two pads. Indeed, the air flow stops and the produced heat, get time to transmit trough the parts. This is simulated by dropping the convection down to this value which is supposed to correspond to stagnant air. Thus, the pads are completely cooled on the outsides which are not contacted with the disc. The FE mesh is generated using three-dimensional tetrahedral elements with 10 nodes (solid 187) for the disc and pads. There are about 185901 nodes and 113367 elements that are used (Fig.16). In this work, a structural analysis will be carried out by coupling thermal analysis. On the figure 22 which represents the distribution after the calculation code of the 3-D model of the temperature - plate disc, yet was chosen corresponding to the maximum temperature Tmax = 346.31 \u00b0 C at t = 1 7271 [s]. The rapid temperature rise of the disk on which two friction tracks lead to an increase of heat stored in the contact area is observed, it is observed that the upper part of the wafer is completely cooled by the effect of convection ambient air" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001347_j.ymssp.2014.06.016-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001347_j.ymssp.2014.06.016-Figure3-1.png", "caption": "Fig. 3. Three-wheeled mobile platform's geometry and kinematics.", "texts": [ " Not widespread solution (however, it can be lastly noted a considerable augmentation of such application) was applied on the control side of the system, where embedded controller cRIO, powered by the LabVIEW, was implemented. For the mechanical part of the design, it is worth indicating the maximum speed, which is achievable while the quality of surveillance system is guaranteed. The value 0.31 m/s was obtained as speed limit. This value was first verified successfully during the virtual tests (bearing in mind mainly capabilities of equipment and study requirements) and subsequently within validation process. An assumed model taken into consideration is presented (Fig. 3). Mobile platform is composed of following main parts: chassis 5, driving system ZN and control system ZK. The driving system consists of two wheels 1 and 2, which thanks to differential mechanism are driven by one electric motor. Wheels rotate about their axes, whose positions are unchangeable relatively to the chassis. Components of the control system are: wheel 3 embedded in steering unit 6, which is driven by the other electric motor 4. Coordinates \u03b11, \u03b12 and \u03b13 are the rotation angles of three mobile platformwheels 1, 2 and 3 respectively", " Solutions were implemented in the LabVIEW, where the detailed survey of multipliers was carried out. The changes in time of dry friction forces were verified only for the \u201csine\u201d type trajectory as the mostly demanding while the mobile platform is conveying. There are demonstrated graphical solutions of the Lagrange multipliers (Fig. 7). Total values of the friction forces TN and TM, acting on appropriate wheels respectively, i.e. on steering wheel and abstractive substitutive wheel used for computation modelling of mobile platform (Fig. 3) can be determined using formula TN \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u03bb21\u00fe\u03bb22 q ; TM \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffi \u03bb23\u00fe\u03bb24 q : \u00f020\u00de However, the limit forces for several wheels are defined by relationships: TN max \u00bc \u03bcN U \u00f0N1\u00feN2\u00de; TMmax \u00bc \u03bcN UN3: \u00f021\u00de where mN, mM\u2014dry friction coefficients for steering and substitutive wheel respectively, N1, N2 and N3\u2014normal reaction forces for wheel 1, 2 and 3. On the basis of the instantaneous values of the Lagrange multipliers, by using Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure22-1.png", "caption": "Figure 22. Distribution of the disk temperature and pads at time t = 1.7271 [s]", "texts": [ " Indeed, the air flow stops and the produced heat, get time to transmit trough the parts. This is simulated by dropping the convection down to this value which is supposed to correspond to stagnant air. Thus, the pads are completely cooled on the outsides which are not contacted with the disc. The FE mesh is generated using three-dimensional tetrahedral elements with 10 nodes (solid 187) for the disc and pads. There are about 185901 nodes and 113367 elements that are used (Fig.16). In this work, a structural analysis will be carried out by coupling thermal analysis. On the figure 22 which represents the distribution after the calculation code of the 3-D model of the temperature - plate disc, yet was chosen corresponding to the maximum temperature Tmax = 346.31 \u00b0 C at t = 1 7271 [s]. The rapid temperature rise of the disk on which two friction tracks lead to an increase of heat stored in the contact area is observed, it is observed that the upper part of the wafer is completely cooled by the effect of convection ambient air. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000139_iet-est.2018.5005-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000139_iet-est.2018.5005-Figure4-1.png", "caption": "Fig. 4 Vectorial representation of machine model at (a), (b) Low speed, (c), (d) High speed", "texts": [ " It is observed from Fig. 3b that Teavg obtained from the method ET2 is always less than Te \u2217 and Teavg may even fall below the lower limit of torque bandwidth (\u0394Te/2) at higher speeds. In addition, from (45), it is observed that by changing the angle \u03b3, torque can be controlled. Figs. 4a and d show the effect of null and AVVs on the variation of angle \u03b3 under different operating conditions. At low speed region as depicted in Figs. 4a and b the change in \u0394\u03b8s occurs with the selection of AVV (see Fig. 4a) and results in a rapid increase in torque due to change in angle \u03b3. With the assumption that the variation in \u0394\u03b8r is low as compared to \u0394\u03b8s, which is a speed dependent. On the other hand, with Nu-VV, the change in angle \u03b3 (see Fig. 4b) occurs due to slow variation of \u0394\u03b8r. Therefore, a slow decrease in torque is observed at low speeds. However, at high speeds, as shown in Figs. 4c and d, \u0394\u03b8r counteracts the change in angle \u03b3 with the AVV and results in a gradual increase in torque. On the other hand, with the Nu-VV, a fast decrease in torque is seen by virtue of fast variation in \u0394\u03b8r. In short, with ESS method ET2, Teavg always falls below the Te \u2217 under all operating conditions. Another torque ESS method (ET3) is defined as follows: dTe(k) = 1, if Mn1 \u2265 Te " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002697_01691864.2020.1810772-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002697_01691864.2020.1810772-Figure2-1.png", "caption": "Figure 2. Error model. (a) Twist error and (b) Initial pose error.", "texts": [ " Note that all of the above vectors in the direction of joint motion and points on the joint axis are specified in the frame S. In [19], gnst0 is also expressed in the matrix exponential form as gnst0 = e\u03be\u0302st , (2) where \u03be\u0302st \u2208 se(3) is the twist of the initial pose transformation. Such, Equation (1) can be rewritten as fst(q) = e\u03be\u03021q1 \u00b7 \u00b7 \u00b7 e\u03be\u0302iqi \u00b7 \u00b7 \u00b7 e\u03be\u0302nqne\u03be\u0302st . (3) In the POE calibrationmethod, the JTCmust be satisfied, but it is a trouble to deal with the JTC directly. Eliminating the JTC in calibration model is a method worth considering; therefore, the Adjoint error model [16] was derived, as shown in Figure 2(a), and the joint twist error can be regarded as a small deviation from the nominal twist to the actual one, then the actual twist can be formulated by the adjoint transformation of e\u03b7\u0302 (see Appendix 2 for details), such that \u03bea = Ad(e\u03b7\u0302)\u03ben, (4) where \u03b7\u0302 \u2208 se(3) is a twist of the error parameter for the adjoint transformation from the nominal joint twist to the actual one, also \u03b7\u0302 can be represented by a sixdimensional vector as \u03b7 = [\u03c9T \u03b7 , \u03bdT\u03b7 ]T \u2208 6. e\u03b7\u0302 can be considered as a rigid body transformation and prescribed by e\u03b7\u0302 = [ R p 0 1 ] , (5) where R = I3 + sin (\u2225\u2225\u03c9\u03b7 \u2225\u2225)\u2225\u2225\u03c9\u03b7 \u2225\u2225 \u03c9\u0302\u03b7 + 1 \u2212 cos (\u2225\u2225\u03c9\u03b7 \u2225\u2225)\u2225\u2225\u03c9\u03b7 \u2225\u22252 \u03c9\u03022 \u03b7 \u2208 SO(3), p = ( I3 + 1 \u2212 cos (\u2225\u2225\u03c9\u03b7 \u2225\u2225)\u2225\u2225\u03c9\u03b7 \u2225\u22252 \u03c9\u0302\u03b7 + \u2225\u2225\u03c9\u03b7 \u2225\u2225\u2212 sin (\u2225\u2225\u03c9\u03b7 \u2225\u2225)\u2225\u2225\u03c9\u03b7 \u2225\u22253 \u03c9\u03022 \u03b7 ) \u03bd\u03b7 \u2208 3; R is represented as a rigid body rotation and p as a rigid body translation", " The actual \u03bea of the revolute joint still satisfies the JTC. (2) For prismatic joint, \u03ben = [ 0 \u03bdn ], Equation (4) will be given by \u03bea = [ 0 Rvn ] . (7) Similarly, \u2016Rvn\u2016 = \u2016vn\u2016 = 1, the actual \u03bea of the prismatic joint also meets the JTC. Hence, the Adjoint model will never violate the JTC. In other words, we do not have to deal with the JTC in our calibration process. Besides, for unifying, the actual value of the initial pose can be given by the left or right action of e\u03b7\u0302st [23], as shown in Figure 2(b). On condition that e\u03b7\u0302st is defined in gnst0 frame, gast0 = e\u03b7\u0302st gnst0 , (8) where \u03b7\u0302st \u2208 se(3) is the twist of the error parameter for the transformation from the nominal pose of end-effector to the actual one. Otherwise, if e\u03b7\u0302st is defined in the current frame, gast0 frame, gast0 = gnst0e \u03b7\u0302st . (9) Themore details of Equation (8) can be consulted in [16], we just consider Equation (9) in this article. Based on Equations (4) and (9), the actual pose of the frame T with respect to the frame S can be written as f ast(q) = eAd(e \u03b7\u03021 )\u03be\u0302n1 q1 \u00b7 \u00b7 \u00b7 eAd(e\u03b7\u0302i )\u03be\u0302ni qi \u00b7 \u00b7 \u00b7 eAd(e\u03b7\u0302n )\u03be\u0302nn qne\u03be\u0302 n st e\u03b7\u0302st " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001137_ipec.2014.6869708-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001137_ipec.2014.6869708-Figure8-1.png", "caption": "Fig . 8. Prototype of proposed BLSRMs. (a) Rotor of 8/10 type. (b) Stator of 8/1 0 type. (c) Rotor of 12114 type. (d) Stator of 12/14 type.", "texts": [ " Accordingly, according to the current error, the switching state S of this winding can be selected as 1, 0, or -1, in which the switching state corresponds to the operating modes of the asymmetric converter. Switching state 1 corresponds to magnetization mode, in which two power switches are excited simultaneously; switching state \u00b0 corresponds to freewheeling mode, in which the winding is short-circuit through an power switch and a diode; and switching state -1 corresponds to demagnetization mode, in which two power switches are turned off and the current flows through two diodes. IV. EXPERIMENTAL RESULTS To verify the proposed control scheme, the tests are executed to the two types. Fig. 8 shows the structures of the test prototypes. And the main parameters of the prototypes are shown in Table II. Based on the control scheme, the experimental system is constructed, as shown in Fig. 9. In the experimental system, the eddy current displacement sensors are used to detect the position of the rotor. The linear range of the sensor is from 0.5mm to 1.5mm, the sensitivity is 5V/mm. Meanwhile, to apply the suspending force load on the shaft, a line is used to connect the shaft and suspending force load" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure8-1.png", "caption": "Fig. 8. (a) Constraint definitions for bolt holes in axial node and bending node; and (b) symmetrical loading conditions.", "texts": [ " To obtain the properties of the printed stainless steel, a series of tensile tests are carried out on six 3D printed dog-bone samples and the test results are shown in Fig. 4. As can be seen, the tested modulus of elasticity, strength and elongation for the 3D printed material vary in different extent. This uncertainty in material properties may cause uncertain results in the structural tests. For numerical simulation purpose, the average of the test data is employed and simplified by a multilinear curve. Fig. 8 shows the averaged result from the tensile tests, and the simplified stress-strain curve used for non-linear simulations. The printing direction of dog-bone sample is also shown in Fig. 5. The design of test rig is evolved during a process in which the main idea of the final design is generated based on the necessities that the final design is capable of testing nodes under various loading conditions, such as bending, shear, axial and combined loads. Fig. 6 shows three test rigs generated during the design process", " 6(d)), the distances between bolt holes and the dimensions of the connecting plates are determined in a modular way in which different loading conditions are obtained by changing the configuration of plates and bolts. Configurations of the test rig for six different loading conditions, including bending, shear, compression, tension, and two combined loading conditions are shown in Fig. 7. To pre-determine the load resisting capacities of the designed structural nodes, non-linear finite element analysis is carried out using Abaqus. To apply loads, the internal surfaces of the bolt holes are constrained to reference points. One set of the constrained joints and their respective reference points are shown in Fig. 8(a). The axial node is simulated under both tension and compression to evaluate the effect of buckling on the node behaviour when it is in compression. The loading condition is simulated by applying displacements or rotations to the reference points. In axial node, displacements in the direction of the connected beam members are applied. In bending node, out-of-plane rotations are applied to the reference points. Fig. 8(b) shows the applied displacements and rotations. The maximum forces applied to each bolt are 5kN and 2.63kN respectively for tension node and compression node, and the maximum bending moment applied to each branch of the bending node is 0.13kNm. The failure modes of tension, compression and bending nodes obtained from numerical models are shown in Figs. 9\u201311. The forcedisplacement curves measured at a reference point of each node are also shown in Figs. 9\u201311. As can be seen, the failure of the tension node occurs in the short members of the node as circled in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002762_j.euromechsol.2020.104125-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002762_j.euromechsol.2020.104125-Figure11-1.png", "caption": "Fig. 11. Change in the signal pattern due to change in defect position in case of inner race defect.", "texts": [ " While computing the defect size, the pattern of the signal when the rolling element passes over the defect, is assumed to be similar to what is shown in Fig. 1(b), showing entry at point B and exit at point C. But, when the rolling element negotiates inner race defect, two types of patterns are observed in the signal because the position of the vibration measuring sensor is fixed, and the position of defect keeps on changing during the rotation of the inner race. Due to this, some impulses are similar to those shown in Fig. 1(b), and some are observed in inverted form. One such instance with a normal pattern of SOI is shown in Fig. 11 (b), while the inverted pattern of SOI is shown in Fig. 11(d). The corresponding position of the defect and sensor location is shown in Fig. 11 (a) and (c), respectively. When it is aimed to determine the defect size automatically, the change in the pattern of the signal may lead to wrong results. Hence a methodology is devised to detect the pattern and reverse the amplitudes if it is as shown in Fig. 11(b). This aids in developing a generalized algorithm for defect estimation. For identifying the pattern change, SOI is extracted, as described in section 2.3.1. For the given SOI, peaks on the positive side and on the negative side are extracted along with their corresponding timestamps, as shown in A.P. Patil et al. European Journal of Mechanics / A Solids 85 (2021) 104125 Fig. 12 (a) and (b). Now, the instantaneous slopes between two adjacent peaks are found out for peaks on the positive side as well as on the negative side" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure5.18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure5.18-1.png", "caption": "Figure 5.18 Solid gap gauge", "texts": [ " NOT GO plug gauges should not enter the hole when applied by hand without using excessive force. These are used to check shafts. Plain ring gauges are ordinarily used only as GO gauges, Fig. 5.16, the use of gap gauges being recommended for the NOT GO gauge. The use of NOT GO gauges is confined to setting pneumatic comparators, and these gauges are identified by a groove around the outside diameter. Plain gap gauges are produced from flat steel plate, suitably hardened, and may be made with a single gap or with both the GO and the NOT GO gaps combined in the one gauge, Fig. 5.18. into a hole or over a shaft while the \u2018NOT GO \u2019 end should not. These are used to check holes and are usually renewable-end types. The gauging member and the handle are manufactured separately, so that only the gauging member need be replaced when worn or damaged or when the workpiece limits are modified. The handle is made of a suitable plastics material which reduces weight and cost and avoids the risk of heat transference. A drift slot or hole is provided near one end of the handle to enable the gauging members to be removed when replacement is necessary" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001045_cdc.2014.7039508-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001045_cdc.2014.7039508-Figure3-1.png", "caption": "Fig. 3. Side view of the vehicles", "texts": [ " In this work, it is assumed that the angles of arrival are available to be directly employed for navigation. Figure 2 summarizes the vehicles coordinates in the horizontal plane of the earth frame. The ground vehicles are parametrized by the coordinates Vi(xi, yi) with xi and yi are constant and i = 1, 2. (x(t), y(t)) are the coordinates of the drone and dgi (t) is the ground distance between the drone and the ith vehicle, i.e. dgi (t) = \u221a (xi \u2212 x(t))2 + (yi \u2212 y(t))2. A side view of the vehicles is shown in Figure 3 where the cross-section plane is the vertical plane containing the UAV and the ith vehicle. This figure shows the height h of the UAV with respect to the ground vehicles as well as the distance di from the UAV to the ith vehicle. The distance di is given as function of the ground distance dgi and height h as d2 i = (dgi )2 + h2. The explicit model for the signals strength Pi (i = 1, 2) shown in Figure 1 is presented in the subsequent section. The propagation model for the signals strength Pi as measured by the UAV\u2019s on-board antenna is given as: Pi = Ti +GTxi +GRx\u2212 PLi (1) where Ti is the transmission power, GTxi and GRx are the gains of the transmitting and receiving antennas respectively, and PLi is the path-loss", " Moreover, in urban macro-cell environment the obstructed or non-line of sight is common for propagation conditions. The building blocks can form either a regular Manhattan type of grid or have more irregular locations. The heights of buildings are over four floors and the density is mostly homogeneous. The NLOS path-loss model is given by [16]: PLi = [44.9\u2212 6.55 log10(h)] log10(di) + 34.46 + 5.83 log10(h) + 23 log10 ( fci 5 ) (3) where h is the height of the base station (the UAV), fci is the transmission frequency and di is the distance as defined in Figure 3. In this paper, the following assumption is made. Assumption 1: The two antennas mounted on the ground vehicles are assumed to be identical and that the transmission powers T1 = T2 = T and the frequencies fc1 = fc2 = fc. Moreover, the antennas are assumed to be perfect omnidirectional and therefore GTxi = 0 and GRx = 0 for i = 1, 2. This assumption is realistic since ground units are generally equipped with similar antennas. It turns out from (1) and (3) that the signals strength can be written as follows: Pi = C1 \u2212 C2 log10(di) ; i = 1, 2 (4) where C1 = T \u221234" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003746_s10409-021-01089-9-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003746_s10409-021-01089-9-Figure9-1.png", "caption": "Fig. 9 Model information of the flexible pendulum", "texts": [ "4 Length A (m) 0.4 Height B (m) 0.1 Maximum thickness tmax (m) 0.012 Minimum thickness tmin (m) 0.01 Width (m) 0.08 Length of the shackle Ls (m) 0.1 Young\u2019s modulus of steel (GPa) 210 Poisson ration of steel 0.3 Young\u2019s modulus of rubber (MPa) 4.41 Poisson ratio of rubber 0.4 1 3 discretized by 1 \u00d7 1 \u00d7 1 , 2 \u00d7 2 \u00d7 2 , 3 \u00d7 3 \u00d7 3 and 4 \u00d7 4 \u00d7 4 element meshes respectively. The simulation time is 1 s. All the other parameters such as the elastic modulus E, the Poisson ratio and the gravity g can be found in Fig.\u00a09. One of the corner points on the free end is chosen to be monitored, which is marked in red in Fig.\u00a09. The varying of the vertical displacement of the observed point is shown Fig.\u00a010. It can be observed that with the increasing of the element number, the displacement curve is getting closer. In other word, the convergence property of the proposed element can be proved. Figure\u00a011 gives the configurations of the 2 \u00d7 2 \u00d7 2 mesh at different moments. The change of the kinetic, potential, elastic and the total energy of the 2 \u00d7 2 \u00d7 2 mesh is presented in Fig.\u00a012. As an isolated system, the total energy remains zero all the time" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003340_tia.2021.3058200-Figure24-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003340_tia.2021.3058200-Figure24-1.png", "caption": "Fig. 24 Demagnetization analysis at 270 elec. deg. rotor position with PM and 3 times of rated current excitation.", "texts": [ " > REPLACE THIS LINE WITH YOUR PAPER IDENTIFICATION NUMBER (DOUBLE-CLICK HERE TO EDIT) < 9 through the other two teeth, which ensures that only little flux goes through the PMs. As shown in Fig. 22(d) and Fig. 23(a), when the rotor positions of the 18/17-pole IFC-BFPMM and 18/13-pole IFC-DSPMM are 270 elec. deg. and 0 elec. deg. respectively, their analyzed PMs have potentially the risk of demagnetization, because their directions of circumferential flux density are opposite to those of PMs. Further, the circumferential flux density distributions of the two machines generated by PM and 3 times of rated current excitation are shown in Fig. 24 and Fig. 25. The used magnet material is N35H. Because their working temperature may reach 120 \u2103, the demagnetization of the PMs will occur with the circumferential flux density lower than 0.35 T. In the 18/13-pole IFC-DSPMM, the area of the analyzed PM where the circumferential flux density is lower than 0.35 T is much smaller, which indicates that this machine has a better anti-demagnetization capability. An 18/13-pole IFC-DSPMM is manufactured and tested to verify the foregoing 2D FE analysis, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002725_tmag.2020.3021644-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002725_tmag.2020.3021644-Figure5-1.png", "caption": "Fig. 5. Flux lines and flux density distributions.", "texts": [ " Downloaded on September 20,2020 at 21:04:18 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 specifications of the discussed DFM-CMG are listed in Table I. As a comparison, CMGs I and II are also considered. The key dimensions and PM volume of the three CMGs are same. The finite-element method (FEM) is used to analyze the CMG performances. Fig. 5 shows the flux lines and flux density distributions of the three CMGs. Compared with CMGs I and II, the air-gap magnetic fields of the DFM-CMG are modulated by the two modulators. Both of the modulators conduct the magnetic flux. From the flux distributions at different \u03b1m-m, it can be found that the IM shift affects the flux paths. The air-gap flux densities and the corresponding harmonic spectra of the DFM-CMG at \u03b1m-m=0 are illustrated in Fig. 6. Fig. 6(a) shows the flux density in air-gap I. The PPN of the fundamental harmonic is 2, which is equal to the PPN of HSR PMs, PH" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003486_j.jmps.2021.104411-Figure22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003486_j.jmps.2021.104411-Figure22-1.png", "caption": "Fig. 22. New design of the end segment for the soft-device testing of the bi-stable flexural-tensegrity beam.", "texts": [ "00N/mm, because the cable is a bit longer than in the hard-device-tested beam due to the new design of the end segments. The prototype lies on a well-lubricated horizontal plane, in order to rule out the effects of the dead weight. The external forces are introduced through wires, redirected towards the vertical by means of pulleys, to which weights have been attached. It is important to assure that the moment applied at the end segments remains constant regardless of their rotation: hence, the end segments have been designed as indicated in Fig. 22. Observe that they are characterized by two appendices: the vertical one is placed in contact with a straight fixed support, well lubricated so to minimize the effects of friction, in order to become a roller constraint; the cable to which the weights are attached is convoluted on the circular appendix with radius 60 mm, centered on the tip of the vertical appendix. When the end segment rotates, the cable unwinds from the circular guide and its tensile force maintains a constant lever arm with respect to the center of rotation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000932_icra.2015.7139531-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000932_icra.2015.7139531-Figure4-1.png", "caption": "Fig. 4: Impact model between the flat end effector and the ball.", "texts": [ " Knowing the state of the ball immediately before and after the hit, one can derive the required end-effector angle and velocity during the hit. To do so the ball velocities are first decomposed into relative normal components v\u22a5,v \u2032 \u22a5 and relative tangential components v\u2016,v \u2032 \u2016 before and after the hit as v\u22a5 = P\u22a5 ([ vx vy ] \u2212 [ vxh vyh ]) , v\u2016 = P\u2016 ([ vx vy ] \u2212 [ vxh vyh ]) , v \u2032 \u22a5 = P\u22a5 ([ v\u2032x v\u2032y ] \u2212 [ vxh vyh ]) , v \u2032 \u2016 = P\u2016 ([ v\u2032x v\u2032y ] \u2212 [ vxh vyh ]) , (17) with the projection matrices P\u22a5,P\u2016 depending on the end effector angle \u03a6 as P\u2016 = [cos(\u03a6h) sin(\u03a6h)] T [cos(\u03a6h) sin(\u03a6h)], P\u22a5 = I\u2212P\u2016. (18) Fig. 4 gives a detailed illustration of the impact model. The hit between ball and end effector is modeled as a partially elastic collision with restitution coefficient \u01eb along the normal direction, resulting in the change of normal impulse \u2206p\u22a5. In tangential direction it is assumed that the change of impulse \u2206p\u2016 during the hit just reduces the relative tangential velocity v\u2016r between ball and end effector to zero. The tangential impulse also leads to an angular momentum L\u2016. Mathematically it can be described by \u2206p\u22a5 = \u2212(1 + \u01eb)p\u22a5, \u2206p\u2016 = \u2212mv\u2016r = \u2212m(v\u2016 + \u03c9 \u00d7 r), (19) \u2206L\u2016 = r\u00d7\u2206p\u2016, with m as the mass of the ball, r as the vector from the center of the ball to the contact point, \u03c9 as the vectorized version of \u03c9 and the operator \u00d7 as the pseudo cross product in 2D" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001651_s00170-016-8545-0-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001651_s00170-016-8545-0-Figure8-1.png", "caption": "Fig. 8 Electrode feeding path", "texts": [ " Shaped electrode design and electrode feeding path planning are the most important steps for EDM. They are interdependent with each other. The design should be modified while it is failed to plan an electrode feeding path or the electrode is too thin. The details of shape electrode design and electrode feeding path planning for shrouded blisks are discussed in another paper [3, 9]. In this paper, the tangent tracking method proposed by Liu and the authors is used for electrode feeding path planning. Figure 8 shows the electrode feed path with multi-axis motion. It is demonstrated by the motion path of the end face centre on the electrode. Machining process was conducted on a six-axis CNC EDM machine tool developed by Shanghai Jiao Tong University. When positive polarity is chosen for Ti6Al4V machining, instability is more likely to occur and it will also cause large wear of electrode than the removal of workpiece, which is very disadvantageous for finishing. While for negative polarity, which was finally used, discharge is more stable and it is easier to obtain better surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003340_tia.2021.3058200-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003340_tia.2021.3058200-Figure5-1.png", "caption": "Fig. 5 Optimized machine topologies. (a) 18/17-pole IFC-BFPMM. (b) 18/13- pole IFC-DSPMM.", "texts": [ " 0093-9994 (c) 2021 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. > REPLACE THIS LINE WITH YOUR PAPER IDENTIFICATION NUMBER (DOUBLE-CLICK HERE TO EDIT) < 4 where is the resistivity of copper, N is the turns per coil, I is the rated current, la is the active axis length, lend is the length of end-winding, S is the area of stator slot, kf is the packing factor. Fig. 5 shows the topologies of two optimized machines. The main parameters of two optimized machines are listed in TABLE II. Fig. 6 compares electromagnetic torque of the initial and optimized designs of both machines. It can be observed that the electromagnetic torques of the optimized 18/17-pole IFCBFPMM and 18/13-pole IFC-DSPMM significantly boost. To get further insight into the characteristics of the optimized two machines, such performances as open-circuit field distribution, open-circuit phase flux linkage, phase back-EMF, dq-axis inductances, torque capability, efficiency and power factor, cogging torque, unbalanced magnetic force (UMF), and demagnetization of PMs are minutely discussed here" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002328_j.mechmachtheory.2020.103873-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002328_j.mechmachtheory.2020.103873-Figure11-1.png", "caption": "Fig. 11. Bending stress on gear flank.", "texts": [ " The bending stress for pinion and wheel are calculated by quasi-static analysis method. The wheel flank was selected to show Mises stress and contact pressure results to save the post-processing. Case 1: Face-milled hypoid gear The Mises stress for wheel in two meshing moments are demonstrated in Fig. 9 . The maximum point is in the center point of contact ellipse. The 2D and 3D diagrams of contact stress are shown in Fig. 10 . The maximum value is in the center of gear flank. The bending stress on pinion and wheel flanks are depicted in Fig. 11 (a) and (b). The positive maximum value is in the root of gear flank. The negative maximum value is in the working part of gear flank. Those stress values represent the tension-compression stress in meshing process. The root stress results of pinion and wheel are shown in Figs. 12 and 13 . The blue line is enveloped from black line in Figs. 12 ( a ) and 13 ( a ). Case 2: Face-hobbed hypoid gear The Mises stress for wheel in two meshing moments are demonstrated in Fig. 14 . The 2D and 3D diagrams of contact stress are shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000267_1.g003407-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000267_1.g003407-Figure3-1.png", "caption": "Fig. 3 Speed vectors, axis systems and wind angles.", "texts": [ " For this purpose, reference is made to the following vector relation of the equations of motion of a soaring vehicle in an atmosphere with moving air (with point mass dynamics considered as appropriate): m R L D mg (1) This relation is expanded to yield the equations of motion in component form for a reference system related to the inertial speed and another one related to the airspeed. For the inertial speed, the speed relative to the Earth is used: Vinert _R (2) Applying Vinert, Eq. (1) can be expanded to yield m _V inert L sin \u03b1w \u2212D cos \u03b1w cos \u03b2w \u2212mg sin \u03b3 mV inert cos \u03b3 \u22c5 _\u03c7 aL1L \u2212 aD1D mV inert _\u03b3 aL2L \u2212 aD2D \u2212mg cos \u03b3 (3) The reference system used in Eq. (3) and graphically addressed in Fig. 3 shows as a constituting feature that the xin axis is aligned with Vinert, with the result that the xin axis is tangential to the flight path with path angle \u03b3. Accordingly, this system is termed path-related reference system in the following. The quantities \u03b1w and \u03b2w are wind-related angles that describe the spatial orientation of the airspeed vector relative to the inertial speed vector [28]. These angles and their relationships with the speed vectors and axis systems are presented in Fig. 3 (with a detailed description given in Appendix A). The reason why the wind-related angles \u03b1w and \u03b2w are used instead of other angle relationships describing the orientation of the airspeed vector relative to the inertial speed vector is that, in this way, the wind effect on kinetic energy can be mathematically described in a straightforward and illustrative manner. This is because \u03b1w and \u03b2w yield rather simple expressions of theL andD components (in terms ofL sin \u03b1w andD cos \u03b1w cos \u03b2w) Fig. 2 Dynamic soaring cycle of spiral or oval type", " In an assessment of the energy characteristics including the total, kinetic, and potential energy, it is shown that the concepts of kinetic energy related to airspeed and to inertial speed are equivalent. There is equality of energy harvesting in terms of the energy gain from the wind which is necessary for compensating for the irreversible energy loss due to the drag work so that energy-neutral dynamic soaring is possible. Thus, it can be concluded that findings and results obtained with one or the other kinetic energy concept are of equal value. In Fig. 3, the speed vectorsVinert,Va, andVw, the path-related axis system xin; yin; zin , the airspeed-related axis system xa; ya; za and the wind angles \u03b1w, \u03b2w, and \u03bcw are shown, with reference made to [28]. The inertial speed vector Vinert is constitutive for the path-related axis system such that the xin axis is aligned with Vinert. The path angle \u03b3 shows the inclination of the xin axis relative to the horizontal plane. The yin axis perpendicular to the xin axis is in the horizontal plane, and the zin axis is perpendicular to the xin-yin plane. For the axis system related to the airspeed, the airspeed vector Va is constitutive to the effect that the xa axis is aligned with Va. The ya axis is perpendicular to the plane that is made up by the xa axis and the lift vector, and the za axis is perpendicular to the xa-ya plane. The wind angle \u03b1w is in the xin-za plane, and \u03b2w is in the xa-ya plane. There is a line of intersection of both planes, indicated by the notation \u201cIL\u201d in Fig. 3. The angular difference between that intersection line and Vinert determines the magnitude of \u03b1w. The magnitude of \u03b2w is determined by the angular difference between the intersection line and Va. The matrix for transforming quantities from the xa; ya; za system into the xin; yin; zin system is given by Min;a 2 664 1 0 0 0 cos \u03bcw \u2212 sin \u03bcw 0 sin \u03bcw cos \u03bcw 3 775 2 664 cos \u03b1w 0 \u2212 sin \u03b1w 0 1 0 sin \u03b1w 0 cos \u03b1w 3 775 \u00d7 2 664 cos \u03b2w sin \u03b2w 0 \u2212 sin \u03b2w cos \u03b2w 0 0 0 1 3 775 (A1) From this matrix, the coefficients aD1;2 and aL1;2 used as abbreviations of trigonometric relationhips in Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001031_j.jbiosc.2014.04.021-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001031_j.jbiosc.2014.04.021-Figure1-1.png", "caption": "FIG. 1. A schematic of the modelled anodic mechanisms where the cytoplasmic NADH/ NAD\u00fe cycle as the electron supplier. Microbes take up substrates (i.e., acetate and/or light) generating carbon dioxide and proton. This process yields electrons for metabolic benefit, i.e., growth, and reduces Medox in the cytosol into Medred. Medred diffuses into contact with the electrode, where Medred reduces the electrode generating electrical current. The oxidized form, Medox, diffuses back through anolyte for reuse by the microbes. Radial red circle highlights the redox cycle supplying the electron to the anode, which is investigated in the present modelling. The iron/sulfur proteins such as ferredoxin (F) and quinone pool (Q) formed in the photophosphorylation and oxidative phosphorylation may also be a potential electron source to be targeted for current production. These reducing species are located in the membrane-surrounded organelles and are not studied in the present modelling. Other cellular compartments, such as flagellum and eyespot etc., are included in the genome-scale network model, but NADH contained in these compartments is not investigated since no evidence indicates that the mediators could reach the reducing power inside these organelles.", "texts": [ " The interactions with an electrode were captured by introducing three reactions into the model reconstruction (Table S1). These reactions represent the net reaction between the reducing equivalents and the electrodes in MFCs, with mediators not shown because they act as intermediates only. Introduction of these reactions creates an additional escape channel for electrons, and their fluxes are subject to the mass balance rule in the FBA modelling. The process of conveying electrons towards an MFC anode is schematically shown in Fig. 1. Multi-objective formulation The extraction of electric current in an MFC creates a new metabolic environment, in which the cell needs to produce excess current carriers in addition to its needs for normal growth. To judge the maximum capabilities to convert low energetic co-substrates (NAD) to their highly energetic counterparts (NADH), the FBA optimization needs to represent a range of priorities for these competing metabolic demands and allows us to evaluate the impact of the enhanced current extraction on cellular metabolism" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003364_012142-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003364_012142-Figure1-1.png", "caption": "Figure 1.Block diagram representation of electrical vehicle system", "texts": [ " In addition to reduced pollution in the environment, electric vehicle gives good performance in terms of its efficiency & torque [1]. The only disadvantage of EV is its cost [2]. Due to the environment concern and less fuel consumption EV are attractive than conventional IC engines [3].For better efficiency of EV, it is necessary to choose motor drive and its control techniques properly. This means electric motor drive system is very much important and it is like heart of the entire electric vehicle system. The following figure 1 represents the block diagram of various parts of electric vehicle system [4]. In the above fig (1), controller part, power converter part and electric motor part represents core of the drive of the EV. It consists of electric vehicle control system and battery management system works together to reduce power consumption [1, 5]. The motor that we choose must have following basic requirements which are [1]. IVC RAISE 2020 IOP Conf. Series: Materials Science and Engineering 1055 (2021) 012142 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001651_s00170-016-8545-0-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001651_s00170-016-8545-0-Figure4-1.png", "caption": "Fig. 4 Blisk blank formed by EBM", "texts": [ " Then the relationship between the blades and the electrode can be recognized from these photos. Therefore, visual identification is very convenient to extract features from the near-net formed blisk. Figure 3 shows the design of the blisk EBM model. The theoretical EDM machining allowance of every single surface on blisk is set as 0.5 mm, which is a little larger than the EBM machining error 0.3 mm. This ensures the EBM blisk blank envelops the theoretical profile of the final blisk even if the maximum error of EBM forming happens. Figure 4 presents the blisk blank formed by EBM. Visual identification is conducted when blisk cutting is finished. At that time, the axial and radial machining references of the blisk have been machined. After the blisk is placed on the EDM machine tool, two coordinate systems need to be set up for visual identification. One is the world coordinate system (WCS) which is the same as the EDM machining coordinate system, and the other is the visual coordinate system(VCS). In the VCS, the main optical axis of the camera is set as the N-axis, and the horizontal and vertical of the camera are U-axis and V-axis respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001201_amr.1038.29-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001201_amr.1038.29-Figure1-1.png", "caption": "Fig. 1: Overview of Fused Deposition Modeling (FDM) process", "texts": [ " Stratasys patented fused deposition modeling, FDM, is the leading extrusion based technique which deposits layers of thermoplastic material to build three-dimensional parts. This process works by converting a thermoplastic material to a monofilament feedstock. This feedstock is fed through the system up to the extrusion nozzles. The material is then heated to a less viscous state by the headed extrusion nozzle and deposited onto the build platform in a precise toolpath that has been determined by the Stratasys software [2]. Fig. 1 depicts the graphic representation of this process. All rights reserved. No part of contents of this paper may be reproduced or transmitted in any form or by any means without the written permission of Trans Tech Publications, www.ttp.net. (ID: 128.210.126.199, Purdue University Libraries, West Lafayette, USA-09/07/15,19:24:55) To allow for the processing and building of complex and intricate geometries a support material is utilized. Each extruded layer that is deposited can consist of model, support material or both, depending on the overall part geometry" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002574_j.mechatronics.2020.102399-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002574_j.mechatronics.2020.102399-Figure3-1.png", "caption": "Fig. 3. Options for set-based collision avoidance of a long, thin obstacle (red). A single, spherical set-based task as defined in (6) covers the entire obstacle, but also includes a lot of space that could be safely accessed by the robot. Several small, spherical obstacles stacked together covers the obstacle much better, but results in a more complex control system than necessary. A single, column-shaped set-based task completely covers the obstacle without also including safe space around the obstacle. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)", "texts": [ " This is derived through the differentiating (6) w.r.t. time. In particular, the Jacobian relates changes in the joints \u02d9 q to change in the distance \u02d9 \u03c3a and depends on the current configuration q . \u02d9 \u03c3a = J a \u0307 q = \u2212 (p o \u2212 p) T || p o \u2212 p|| J \u02d9 q (7) ere, p denotes the position of the end effector and J is the corre- ponding position Jacobian (2) . .1. Column-shaped obstacles Although general, a spherical shape of an obstacle as (6) is not ecessarily the optimal representation for any obstacle, as illusrated in Fig. 3 . For example, a spherical shape around a long, hin object would be highly conservative as a lot of safe space ould be excluded from the workspace. Another option is to define ultiple small spheres to avoid defining available space as non- ccessible, but this would lead to a more complex control system ith multiple set-based tasks, which, although feasible, is not as ptimal as defining a set-based task based on a more fitting shape. he kinematics and implementation to include column-shaped obtacles in the set-based framework for redundant robotic systems re presented below" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000959_s00502-014-0272-3-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000959_s00502-014-0272-3-Figure1-1.png", "caption": "Fig. 1. Structures of the radial and axial flux PM machine with surface mounted magnets [3]: (a) two pole radial flux machine with inner stator and distributed stator winding; (b) single sided eight-pole axial flux machine with stator tooth coil winding", "texts": [ " The additional motor mass inside the wheel increases the unsprung mass of the vehicle and leads to a higher stress of the wheel suspension [1, 4]. Therefore the motor should be as light as possible; hence a high power density is needed. Due to this, high energy permanent magnet technology with a high pole count is used to reduce the mass of the motor. Two different structures of the permanent magnet (PM) synchronous motor are investigated in this report: the axial flux machine (AFM) and the radial flux machine (RFM) (Fig. 1). The radial flux structure is the commonly used structure of the permanent magnet synchronous machines. The axial flux machine is a special type of the permanent magnet synchronous machine and is advantageous especially when a short axial length of the electric motor is required. Figure 1 shows the two different directions of the magnetic field lines through the air-gap of the machines. In case of the radial flux machine the magnetic field lines pass radially and in case of the axial flux machine the field lines pass axially through the air-gap. The different directions of the flux lead to different coil arrangements inside the machines and therefore to alternative iron structures. The radial flux machine has a cylindrical iron structure, while the axial flux machine has a disc-like motor shape. Due to this, the axial flux machine has a short axial length and is especially suitable for wheel hub drives. Different coil and rotor arrangements for the axial flux machine are possible (single-sided, double-sided and multi-stage machines), which are summarized in Fig. 2. The single-sided machines (Fig. 1b) have one stator and one rotor and exhibit a high unbalanced axial magnetic pull between the stator and the rotor, resulting in a high axial force putting high requirements on the bearing design. The multi-stage machines are a stacked combination of the double-sided machine type, and allow a torque scaling at the expense of an increased axial length. The double-sided machines fit best to the requirements of a wheel hub drive. These double-sided machines can be divided into two categories: Motors with internal rotor (AFIR: axial flux internal rotor) and motors with internal stator (AFIS: axial flux internal stator)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001720_ecce.2014.6953594-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001720_ecce.2014.6953594-Figure2-1.png", "caption": "Figure 2. (A) Schematic of a part of the stator (iron part, slots, and windings), (B) Centerline cut including one slot and windings, (C) The winding which is used in the 2D finite difference thermal model", "texts": [], "surrounding_texts": [ "The analysis presented above is quantified using a twodimensional steady state finite difference (FD) model of a copper coil from central part of the stator to the end winding of an electric machine as illustrated in Figs. 2 and 3. In this analysis, the convective heat transfer coefficients are prescribed. Detailed numerical simulations, described in the next section, are required to obtain the actual convective heat transfer coefficients. As it is shown in Fig. 3, the thermal network includes the insulation conduction thermal resistance inside the stack ( , , ), the insulation conduction thermal resistance through the iron yoke inside the stack ( , ), the insulation conduction thermal resistance, and the convection thermal resistance at the end winding ( , , , Eq. (8)), the copper axial conduction thermal resistance through the winding ( , ), and the Ohmic power loss per volume of the winding conductor ( \"). Thermal resistances and the power loss are approximated for each node in the thermal network based on the known coil, yoke and insulation geometry and characteristics, and also the system current density. and are the temperature of the outer surface of stack and the ambient temperature around the end winding which are assumed constant (40 \u00b0C) in this simplified analysis. This configuration is similar to, but not the same as, water jacket cooling at the stator outer surface. Water jacket cooling has been effectively used in compact Azipod (ABB, Inc) electric propulsion units and also in many mine motors [9]. A principle difficulty with water jacket cooling is limited cooling of the inner regions of the machine, which is largely taken care of by FF cooling concept presented here. Thermal resistances are defined as: , , , , (10) , , (11) , , (12) , , (13) where , , , , , , , , , , , and are the stack winding insulation thickness, the associated stack winding and insulation surface, the stator yoke thickness, the stator yoke thermal conductivity, the winding length, the winding cross-section, the resistivity of copper, and the current density, respectively. This model was applied for 2, 4, and 6 poles (P), 25 MW electric machine. Details of these electric machines are presented in Table 1 (data provided by ABB Inc.). Table 1. Electric machines data used in 2D thermal model Results from the 2-D FD thermal model are presented in Fig. 4. In these figures the solid blue lines show the temperature distribution of the winding section inside the stack, while the red dashed lines show the temperature of the coil in the end winding section (Fig. 3). As shown in Fig. 4 (A) for 25 MW (class F) air cooled machines with average convection coefficient of 50 W/m2-K [7] on the end winding surface, the maximum temperature is approximately 145 \u00b0C which occurs in the end winding segment. For this case, heat transfer in the stack is due mainly to radial heat flow through the cooling paths in the iron stack. Fig. 4 (B) shows the temperature distribution if ferrofluid (FF) cooling is used which allows for a reduction in main wall insulation thickness around the end winding section (i.e. using the dielectric property of the FF), but with the same average convection coefficient as used for the air cooling (case A) - which is acknowledged to be low for liquid cooling. Our analysis shows that the maximum temperature for this case is reduced by approximately 15 \u00b0C, while the end winding temperatures are reduce by approximately 30 \u00b0C. Also the temperature distribution in the winding (in both stack and end winding) is changed. Fig. 4(C) shows the temperature profile along the winding for a FF cooled system with h=1000 W/m2-K (expected from FF cooling), where the insulation thickness of coils in the end winding section is reduced to 0.1 mm (i.e. the turn insulation only). In this case the maximum temperature of the 6 pole winding is reduced by about 45 \u00b0C compared to the air cooled machine (case (A)). Also the centerline temperature and the average temperature of the windings in the stack are reduced. Fig. 4(D) shows the case where the electric current density is increased by ~45% (for the 2 pole machine, but can be increased further for the 4 and 6 pole machines while keeping Tmax < 150 \u00b0C) using the same h and tin,ew parameters used in case (C). Due to the larger air gap diameters of the higher pole number machines, the stack lengths of these higher pole number machines is shorter than those for the lower pole number machines. As such, there is effectiveness of this novel cooling technique for higher pole number machines. Increasing the electric current density (electrical loading) by providing efficient FF cooling on the end windings is promising for increasing the power density not only due to directly increasing the electric loading, but also due to the possible elimination of the air passages within the air cooled machine and the reduction of the stack length of the machine. For example, using equations presented in section A the noted 45% increasing in electrical loading (current density) with efficient FF cooling can reduce the air gap diameter by about 20% (or stack length by about 40%), so that active mass/volume could be reduced by about 1/3rd for such these machines. Comparing results in Fig. 4 indicates that all three types of machines approach the same end winding temperature, but that higher pole number machines are more compatible with FF cooling where temperature profiles along their windings shows more uniformity. For example, the 6 pole machine exhibits the lowest temperatures (and also the lowest temperature rise) at the centerline and also along length of the windings with FF cooling. This is also that case even when the current density is increased (Fig. 4(D)). These results can be explained by considering the geometry of these machines where higher pole number machines have shorter stators for a given torque rating. So the losses are conducted through the windings to the end windings and are dissipated in the FF cooling medium instead of being dissipated though a high thermal resistance radial path through the slot insulation and stator. III. NUMERICAL SIMULATIONS OF END WINDING HEAT TRANSFER CHARACTERISTICS Some preliminary proof of concept three dimensional simulations using COMSOL Multiphysics simulation software (Version 4.3b) have been carried out to compare the proposed cooling approach with the more common air cooling approach for cooling end windings of electric machines. These simulations are based on 2 [kW] of heat dissipation per end of an electric machine (96% efficient 100 kW machine (data provided by ABB Inc.)) with prescribed enclosure wall temperatures of 80 \u00b0C (dimensions are shown in Fig. 5). The thermal boundary conditions for ferrofluid cooling (FF filled enclosure) are shown in Fig. 6. It should be noted that in these simulations only heat dissipation and cooling of the end windings are considered. Note that axial symmetry allows the study of a coil segment. The magnetic boundary conditions for modeling the induced magnetic field in the TSFF around the end winding are presented in Fig. 7. For modeling the thermomagnetic convection of temperature sensitive ferrofluid fluids, the fluid is assumed as a single phase Newtonian fluid. This assumption is proper for the description of dynamic phenomena of the magnetic fluid from a macroscopic view [12]. Also, the fluid is electrically non-conducting, thus the displacement current is negligible [12]. The ferrofluid properties are as: \u03c1=1044 kg/m3, Cp=1616 J/kg.K, \u03bc=0.0033 Pa.s, k=0.16 W/m.K, Tc= 358 K, Ms=1\u00d7104 A/m, where Tc and Ms are the ferrofluid Curie temperature (the temperature at which all net magnetization is lost), and the saturation magnetization (the maximum attainable magnetic moment per unit volume of the magnetic fluid), respectively. The governing differential equations for simulating the thermomagnetic convection of the single phase ferrofluid in the presence of magnetic field from the end winding, neglecting gravity effects are [12]: Equation of continuity: \u2207. =0 (14) Equation of motion: \u03c1 (\u2207. ) =-\u2207p+ \u03b7\u22072 + \u03bc0M\u2207H (15) Equation of energy conservation: \u03c1C( .\u2207)T =\u2207.(k\u2207T) (16) where , \u03c1, p, \u03b7, C, \u03bc0, k, M and H are the velocity vector (m/s), density (kg/m3), pressure (Pa), dynamic viscosity (Pa.s), specific heat (J/kg.K), permeability of vacuum (4\u03c0\u00d710-7 N/A2), thermal conductivity (W/m.K), scalar magnitude of the magnetization (A/m), and the scalar magnitude of magnetic field intensity (A/m), respectively. The last term in Eq. (15) is the magnetic body force or Kelvin body force (KBF), which is the body force per unit volume that the fluid experiences in a spatially non-uniform magnetic and thermal field [12,15]. As previously mentioned, the magnetic body force is a function of the FF magnetization (M, which is a function of magnetic field intensity and temperature [12, 15]), and the magnetic field gradient (\u2207H). So this force can be altered using the magnetic properties of the FF (such as saturation magnetization and Curie temperature of the FF), thermal and magnetic fields gradiants. The induced magnetic field inside the ferrofluid which is represented by the magnetic flux density and the magnetic field intensity conform to Maxwell\u2019s relations in static form [12]: \u2207\u00d7 =0 (17) \u2207. =0 (18) \u00b5 (19) Among the properties of the ferrofluid, the most important one is the fluid magnetization, M. For an electrically non-conducting, incompressible fluid this property is a function of the magnetic field intensity and fluid temperature [12,13]. In this study, ferrofluid magnetization is described by the Langevin function [12- 14]. This function can be used for wide ranges of magnetic field intensity and temperature: M=Ms L(\u03be), (20) L(\u03be)=coth\u03be-1/\u03be (21) \u03be= \u03bc0mH/kBT (22) Here, Ms is the saturation magnetization of the magnetic fluid, and L(\u03be) represents the Langevin function. The Langevin parameter \u03be is the ratio of magnetic to thermal energies [12-15]. In this parameter m and kB denote the magnetic moment of magnetic particles in the ferrofluid, and Boltzmann constant, respectively. Temperature results for two cases of cooling the end winding surfaces are presented in Fig. 8. Fig. 8(A) considers thermomagnetic convection effects of a TSFF filled enclosure while Fig. 8(B) considers typical air cooling (from wafters on rotor shaft with equivalent convective heat transfer coefficient as 50 W/m2-K [7]). As Fig. 8(A) illustrates (temperature profile and flow velocity vectors), temperature sensitive magnetic fluid motion occurs due to both the gradients of the magnetic field and the temperature. The magnetic fluid is attracted toward regions with larger field strength, while near the heat source the fluid temperature approaches the Curie temperature of the ferrofluid. In this region the fluid loses its attraction to the magnetic field, and is displaced by colder fluid [12, 15]. This application utilizes FFs whose magnetic properties are strongly influenced by temperature. The TSFFs are chosen such that the material undergoes a substantial drop in magnetization as it approaches the ordinary working temperature of the device to be cooled. As a practical matter, this ordinarily means that the device operating temperature is close to or just below the Curie temperature (the temperature which all magnetization is lost) of the chosen material. A Curie temperature well above or below the device operating temperature will fail to perform in the context of the present application [6]. The presented 3D, steady state, laminar thermomagnetic convection of the TSFF inside the enclosed end winding of the electric machine is established using COMSOL Multiphysics version 4.3b simulation software based on the solution of coupled partial differential equations (PDEs) using the finite element method, and the mesh network includes about 4\u00d7106 elements. This model includes the coupling of three fundamental phenomena, i.e. magnetic, thermal, and fluid dynamic features. In this numerical model, the temperature variation in the temperature sensitive ferrofluid leads to a change in the magnetization and affect the magnetic force acting on the fluid in the presence of the end winding magnetic field. The fluid flow affects the convective heat transfer and the temperature distribution inside the magnetic fluid and this in turn affects the driving force on the fluid. Alternatively, the fluid temperature distribution can induce a magnetization variation within the magnetic fluid and thus may vary the driving force [12]. As these results show, the thermomagnetic TSFF filled cooling (Tmax=87 \u00b0C) reduces the maximum temperature rise (~7 \u00b0C) to about 25% of the air-cooled system (Tmax=107 \u00b0C) temperature rise (~27 \u00b0C). The maximum magnetic field intensity around the windings in this simulation is calculated to be ~20 kA/m. These results show that thermomagnetic convection enhances the cooling of end winding of electric machine due to two effects. The first is due to the higher thermal conductivity of the FF and the second is due to thermomagnetic convection of the TSFF. Lower temperature rise, more uniform temperature distribution, and higher heat dissipation rates are expected when using the proposed thermal management approach for cooling of the end windings of electric machines. Other aspects of this approach are; i.) the system is self-regulating (i.e. as the heat load increases, the magnetization inside the FF drops and the driving force increases, such that the fluid circulates at higher speeds and transports more heat), ii.) the cooling fluid flow is controlled by the existing magnetic field and waste heat, thus no energy consumption or additional devices are needed, iii.) there is no \u201cpore-wick\u201d structure as in capillary pumped loops (i.e. heat pipe), no electrically conductive elements, no moving parts, and no flow pulsation as in reciprocating pump systems; all of which improve the reliability of the system, and iv.) since the property of FF depends on the suspended magnetic particles and the base carrier liquid, there is a large range of possible properties for the FF, allowing it to be tailored towards the needs of the electrical generator application [8]. IV. CONCLUSION In this paper, a series of analytical and numerical studies have been carried out to gain insight into the concept of thermomagnetic convection effect of a temperature sensitive ferrofluid as cooling and dielectric media in electric machines thermal management. Using the analytical model a relation between cooling parameters and power density of electric machines was developed which showed power density of an EM can be altered using the electrical insulation thickness of the end winding and coefficient of convective cooling on the end winding domain. Also a twodimensional steady-state thermal model using the finite difference method was established and applied to 25 MW class (2P, 4P, 6P) generators for different scenarios of the electrical insulation thickness and coefficient of convective cooling. In the end, a simplified numerical simulation using COMSOL Multiphysics simulation software for thermomagnetic convective cooling of the end winding of typical electric machines was presented and compared with air cooling. This work is an introduction to thermomagnetic convective cooling of EMs in order to show the methodology for evaluation of such systems. Future works will include extending the numerical simulations considering effect of gravity body force on the cooling FF circulation in the end winding domain, and also, experimental investigations on dielectric properties of FFs and examination of this approach on a EM end winding experimentally. Our preliminary analysis showed that this cooling technique holds more promise for higher pole number machines, thus the use of this technology may also be suitably used for propulsion motors. Using the proposed cooling technology in the 5-40 MW generator scale is possible as the various loadings and aspect ratios of the generators tend to be quite similar. The effects of rotor magnetic leakage along with the native magnetic field of the end windings in the end regions of the generator for wound field synchronous machines have the possibility of accentuating the turbulence and convection characteristics of the FF flow. Limited rotor losses of PM machines may require only an outer surface waterjacket around the stator, while the significant rotor losses of the wound field or induction type machines would likely necessity a dedicated rotor cooling system, which would further add to the possible benefits of the proposed cooling technology." ] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure17-1.png", "caption": "Figure 17. Variation of the total deformation of the model (full scale)", "texts": [ " It is thus judicious to choose a refined grid, because the solution becomes more exact by increasing the number of nodes of the grid. a) Influence smoothness of the mesh For that, one considered a second type of mesh finer and refined in the friction tracks, (Fig.16) appears. The element used in this mesh is a SOLID 187 and the total time of the simulation is equal to 8331.328 (s). This new mesh (type M2) consists of 11 3367 elements TE with 4 nodes, that is to say 18 5901 nodes. It is thus much finer than the mesh M1, Fig. 15 (c) used up to that point. Figure 17 shows the various configurations of displacements of the order \u00ecm of the model according to time, while keeping the symmetrical form compared to the vertical median plane. The total deformation is reached at the end of braking and varies between 0 to 52,829 \u00ecm . On the model of the not-deformed inner pad one has a degradation of the colors ranging from yellow and green to the red when the critical value is located on the higher radial edge of the deformed pad, shown in figure 15 by red color.This is due to the elastic modulus of of pad which is lower Copyright \u00a9 2015, IGI Global" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003519_s10846-021-01359-5-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003519_s10846-021-01359-5-Figure6-1.png", "caption": "Fig. 6 Custom designed and manufactured follower robot with appropriate sensors for full autonomy", "texts": [ " Therefore, based on the user\u2019s input through the analog sticks, the leader robot moves in the same direction with given velocity using Eq. 2. Four 78 kg-cm rectangular gear box motors are used for this purpose. The follower robot\u2019s main body and wheel assembly are similar to that of the leader robot, the difference being that it is equipped with more sensors and their mounts. In the front acrylic sheet, there are 3D printed mounts for ultrasonic sensors inclined at an angle of 60\u25e6, as shown in Fig. 6. The upper turntable has an additional constrained degree of freedom in a linear direction towards the leader robot\u2019s force. Furthermore, the turntable is coupled with a pair of spur gears in 1:2 ratio to sense the angle at which the RL applies force. Hence, both the magnitude (d) and direction of the force (\u03b8 ) are sensed in this manner, which can be] seen in the assembly displayed in Fig. 7. Figure 8 displays the overall control structure of the follower robot. The ultrasonic sensor is mounted in an inclined fashion so to detect obstacles in the forward-right and forwardleft positions of the robot. Additionally, the upper platform contains a 40mm B50K\u03a9 slide potentiometer, arranged such that the slide is along the direction of force of the leader robot. The turntable is coupled with a 10K\u03a9 rotary potentiometer through a pair of spur gears, as shown in Fig. 6. All these are sensed by an Arduino UNO running at 16MHz. It computes an obstacle-avoidance trajectory and feeds signals to two two-channel L293n motor drivers using Eq. 2. These drivers are powered with a 11.4V Li-Po battery. Four 78 kg-cm rectangular gear box motors are used for this purpose. 6Methodology For a collaborative manipulation of the multi-robot (or two robots) system, the leader RL and follower RF robots are used to carry and transport payload\u201d through an environment with \u201cOi\u201d Objects" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000158_s11041-019-00347-9-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000158_s11041-019-00347-9-Figure5-1.png", "caption": "Fig. 5. Scheme of fabrication of a specimen by laser cladding: a) single track; b ) several tracks; 1 ) deposited track; 2 ) remelting of substrate; 3 ) zone of single remelting; 4 ) zone of double remelting.", "texts": [ " The specimens were pressed in two directions, i.e., along and across the growth direction. Figure 4 presents the microstructure of specimens in the initial condition (without heat treatment). We can see a cast microstructure represented by dendrites with well distin- guishable microscopic pools from each deposited track with a height of 6 \u2013 8 mm. The transverse lap exhibits dendrites with repeated sizes, while the length of the dendrite arms in the longitudinal lap differs. The process of growth of a specimen is presented in Fig. 5. The microstructure formed in a single track is homogeneous. Deposition of the second and subsequent tracks yields two zones, i.e., zones of single and double remelting. The microstructure in such zones differs due to the difference in the cooling conditions. Figure 6 presents the microstructure of specimens after single and double remelting. In the zone of single remelting the dendrite arms are extended to 200 \u2013 300 m, while in the zone of double remelting the maximum length of dendrite arms does not exceed 100 m" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003460_j.ast.2021.106644-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003460_j.ast.2021.106644-Figure1-1.png", "caption": "Fig. 1. System configuration and gimbal frame.", "texts": [ " The attitude dynamics of the spacecraft utilizing a VSCMG is described here. In particular, the total angular momentum of the spacecraft can be expressed as h = J\u03c9 + Icg \u03b3\u0307 g\u0302 + I ws s\u0302 (7) where is the wheel angular velocity and \u03b3 the gimbal angle. s\u0302, t\u0302 and g\u0302 is the triad of unit vectors constituting the gimbal frame G that is attached to VSCMG. Note that the gimbal angle relates the gimbal frame to the body one through the following rotation matrix:\u23a1 \u23a3 s\u0302 t\u0302 g\u0302 \u23a4 \u23a6 = \u23a1 \u23a3 cos\u03b3 sin\u03b3 0 \u2212 sin\u03b3 cos\u03b3 0 0 0 1 \u23a4 \u23a6 \u23a1 \u23a2\u23a3 b\u03021 b\u03022 b\u03023 \u23a4 \u23a5\u23a6 (8) In particular, Fig. 1 shows the two frames. The attitude dynamics are described by: J0\u03c9\u0307 = J0\u03c9 \u00d7 \u03c9 + u + d (9) where \u03c9 is expressed in body fixed co-ordinates and: J0 = J \u2212 J (10) where J is the real inertia matrix, J0 is the assumed constant matrix and J is the uncertain, time-varying, inertia induced by the VSCMG. u is: u = \u2212I ws\u0307s\u0302 \u2212 I ws \u03b3\u0307 t\u0302 (11) and d: d = Icg \u03b3\u0307\u03c9 \u00d7 g\u0302 \u2212 I ws \u03c9 \u00d7 s\u0302 + d2 (12) where d2 the torque induced uncertainties in the inertia of the spacecraft attributed to the VSCMG, thruster off-set and the spacecraft environment (which we assume in LEO to be solar radiation pressure torque, magnetic torque due to parasitic current and gravity gradient)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001534_s00170-015-8269-6-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001534_s00170-015-8269-6-Figure7-1.png", "caption": "Fig. 7 Schematic of the photographing of the interaction between powder particles and molten pool", "texts": [ " In order to verify the opinion, an ingenious experiment was designed for the observation of the interaction between powder particle and molten pool. The observation system is composed of a high-speed camera, a zoom lens, an auxiliary source, and software. In order to eliminate the influence of the particles rebounded by the substrate in the observation, a single-layer deposition on a sheet with only 1 mm thickness was performed and a quartz glass sheet was used to protect the camera. The schematic of the observation experiment is shown in Fig. 7. Figure 8 shows the photos captured by using the highspeed camera system with 500 \u03bcs exposure time and the interaction between the feed powder and the molten pool. The thin white lines in Fig. 8 are the moving trajectories of the powder particles in the exposure time. It can be found that one specified particle was fed into the rear of the molten pool and then rebounded from the position, which indicates that the powder particles are rebounded from the shallow part of the molten pool edges. It can also be found that the white line of the rebounded particle is shorter than that of the incident particle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000491_aim.2019.8868892-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000491_aim.2019.8868892-Figure2-1.png", "caption": "Fig. 2. Structure of the mobile control system", "texts": [ " The operating range and angle of view is 0.5m to 20m and 270 degrees respectively, having a high update rate of 50 Hz. The Kinect camera is connected to the laptop placed on the mobile system to get the depth image and helps in generating a 3D cloud map of the environment. Lastly, a microphone is connected to a static computer (PC1), which communicates with the mobile computer (PC2) placed on the mobile robot via secure shell (SSH). The hardware connection of the mobile voice control system is shown in Figure 2. The voice control module is composed of different modules, like automatic speech recognition (ASR) module and the control module. ASR system examines the speaker\u2019s voice and communicates to the robot to perform the action according to the user\u2019s order. The control module comprehends the speech command and executes the action accordingly. The flow of command system is described in Figure 3. The system of voice command is complicated, when a user commands to the robot, the module detects the user\u2019s voice by extracting features from it" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002858_s12555-019-0975-7-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002858_s12555-019-0975-7-Figure2-1.png", "caption": "Fig. 2. Virtual prototype of 3 DOF wearable exoskeleton.", "texts": [ " To fit in with the upper limb of the operator, 3 DOF mechanical structure is designed as shown in Fig. 1, with two rotational DOFs at the shoulder, one rotational DOF at the elbow. Moving requirements will be guaranteed by the cooperation of the DOFs. Virtual prototyping technique is a method in the process of product development before physical prototype. Mechanical structure of wearable exoskeleton is highly anthropomorphic and is attached to human body. The structure assignment and virtual prototype is shown in Fig. 2. It is done in SolidWorks by creating every generated geometrical part and combing into an assembly to debug mechanical motions, fit and function as well. The virtual prototype simulates the movements for recovery training: shoulder flexion/extension, elbow flexion/extension and the whole limb longitudinal rotation. Hence, so designed virtual prototype is adopted as the controlled object to evaluate the performance of control scheme. 2.2. Ultra-local model In general, dynamic model of the wearable exoskeleton can be described by Lagrangian dynamic equation in the following form: M(q)q\u0308+C(q, q\u0307)q\u0307+G(q)+ \u03c4d = \u03c4, (1) where q, q\u0307, q\u0308 denote the joint angle, joint velocity, joint acceleration of the system, M(q) denotes the inertia matrix which is symmetrical and positive definite, C(q, q\u0307) denotes the centrifugal and Coriolis term, G(q) denotes the gravitational term, \u03c4d is the unknown disturbance, and \u03c4 denotes the torque acts on each DOF" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001601_s00170-016-8487-6-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001601_s00170-016-8487-6-Figure1-1.png", "caption": "Fig. 1 The schematic presentation of Exp-ECAE die together with its geometrical parameters", "texts": [ " In this research work, the product quality (i.e., the magnitude and the homogeneity of imposed plastic strain) of expansion equal channel angular extrusion (Exp-ECAE) process is investigated. After designing and conducting the simulated experiments, the results are gathered to train a feed-forward ANN. This ANN is capable of estimating the magnitude and homogeneity of imposed plastic strain. Finally, by linking a GA to the trained ANN, optimum values of geometrical parameters are determined for Exp-ECAE process. As Fig. 1 illustrates, an Exp-ECAE die simply involves a couple of perpendicular cylindrical channels which intersect at a spherical cavity. The centerlines of channels are located next to the center of spherical hollow with a distance of e=2 mm (Fig. 1). At the beginning of process, the punch pushes the billet with an initial diameter of D0 until the billet reaches the ballshaped region. Then, the bar gradually expands and fills the hollow. Afterwards, the metal flows from the cavity to the exit channel and is, therefore, extruded to its initial diameter [13]. When the billet is placed in the inlet channel (Fig. 2a), by the punch movement, the specimen reaches the spherical hollow and, due to the eccentricity of e, touches the die wall at a single point", " Finally, by superimposing two recent components, the total effective strain attainable by means of an Exp-ECAE can be formulated as: D0\u03b5Total \u00bc \u03b51 \u00fe \u03b52 \u00bc 4ln D1 . D0 \u00fe 2cot \u03d5 . 2\u00fe \u03c8 . 2 \u00fe \u03c8cosec \u03d5 . 2\u00fe \u03c8 . 2 h i. ffiffiffi 3 p \u00f04\u00de Experimental tests were performed using the AA6063 aluminum alloy [14]. Rods with 20 mm in diameter were purchased then machined into the billets with a diameter and length of 15 and 120 mm, respectively. The billets were initially annealed for 2.5 h at 700 K. An Exp-ECAE die with perpendicular channels (\u03d5 =90\u00b0) was used for experiments. According to Fig. 1, geometrical parameters of the die including D0, D1, r, and e were respectively measured to be 15, 23, 1, and 2 mm. Figure 4a exhibits the unassembled halves of die. The ExpECAE die together with the tool accessories are illustrated in Fig. 4b. In order to reduce the interfacial friction, MoS2 was used as lubricant. Several samples were processed through the Exp-ECAE at different temperatures and with various ram velocities while the load-displacement variations were recorded during the experiments", " Preliminary FE analyses revealed certain inhomogeneous distributions of the effective strain within the cross section of Exp-ECAE products. This fact causes a heterogeneous distribution of mechanical properties in the product. Therefore, the optimization of die geometry was conducted to induce the plastic strain as large and as homogeneous as possible into the sample. To provide an overall view about the optimization process, Fig. 5 depicts the flow chart of the applied approach. Geometrical parameters (as shown in Fig. 1), including sphere diameter (D1), channel diameter (D0), fillet radius at the intersection of the sphere and the channels (r), and eccentricity of the ball cavity center with respect to the centerlines of the channels (e) were nondimensionalized then involved in the Exp-ECAE optimization. These parameters are also presented in Fig. 6a. As this figure shows: 0\u2264R0 \u00fe r \u00fe e\u2264R1 \u00f05\u00de According to Fig. 6a due to unclear position of the fillet radius center, Eq. 5 is a rough description. Dividing this equation by R1 yields: 0\u2264R0 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000588_9783527822201.ch15-Figure15.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000588_9783527822201.ch15-Figure15.7-1.png", "caption": "Figure 15.7 Possible strategies to grow straight or bend: (a) a symmetric deposition of material results in a straight growth, continuous deposition; (b) a bending obtained by the deposition of the same amount of layers with different thickness; and (c) a bending obtained by the deposition of a different amount of layers with the same height, reversing deposition.", "texts": [ " By tuning the heater temperature, an optimal structure strength is obtained in both air and soil. The material plays an important role for an effective penetration of the soil. It needs to be externally sticky enough to permit the layers to attach, and at the same time needs to be semisolid in the internal part, in order to be strong enough to overcome soil pressure and push the tip ahead. When the material is deposited uniformly along the plotting circumference, the resulting structure is straight (Figure 15.7a). Adding a different amount of material on opposite sides of the robotic root structure creates an asymmetry that results in a bending in the lower deposition direction. In particular, the bending can be obtained by two methods: (A) Continuous deposition: A deposition of the same number of layers k with different thicknesses h1 and h2 (Figure 15.7b) (B) Reversing deposition: A deposition of different number of layers n and m with the same thickness h (Figure 15.7c). Continuous Deposition (Method A) With this deposition strategy, the direction of the rotation is never inverted and the bending is obtained by tuning the layer height along the deposition circumference. A curvature (Figure 15.8) can be obtained by setting properly the layer thickness h at a generic point P \ud835\udefc along the circumference (with \ud835\udefc angle of P with x-axis): P \ud835\udefc = ( d 2 cos \ud835\udefc, d 2 sin \ud835\udefc ) (15.9) Considering \ud835\udf03 the growing direction with respect to the x-axis, in that position the deposition should be minimum (height h1) and in the opposite position (\ud835\udf03+\u03c0) the deposition should be maximum (height h2)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000549_0731684419888588-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000549_0731684419888588-Figure2-1.png", "caption": "Figure 2. 3D compression-twist cellular structure.", "texts": [ "16 On the other hand, tetra-chiral and hex-chiral have stable Poisson\u2019s ratio, which is close to 1.15 For effective Young\u2019s modulus, its value could increase with the increase of cell number,17 and for out-of-plane elastic properties, chiral structures with more ligaments have better transverse shear modulus. Therefore, tetra-chiral cell not only has large Young\u2019s modulus and stable NPR but also is easy to make up 3D cellular structure. The 3D CTCSs are repeated in three directions, which are shown in Figure 2. The 3D cell is made of the 2D chiral cell, which looks like a cube. Every surface of the cube is a 2D cell, and thus the cell walls of the 2D cell are connected. Therefore, the mechanical properties are determined by the shape of the 2D cell. The mechanical properties of cellular structures are related to the relative density, so it is very important to work on the relative density before studying the axial compression performance of the CTCS structure. The structure of the 2D cell is shown in Figure 3(a), and t is the thickness of the cell wall, l is the length of the cell, and h is the angle between the cell wall and the cell diagonal" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001090_0954409714548101-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001090_0954409714548101-Figure1-1.png", "caption": "Figure 1. Heat flux models for calculating temperatures due to a moving heat source.", "texts": [ " The surface temperature due to a certain heat input can be computed using a model based on a one-dimensional heat flux flowing into the body. Let us assume that the source travels and heats the surface faster than the heat of previous inputs can reach the same position using the conduction process. The temperature patterns due to one-dimensional flow into the body up to a certain point in time can be arranged along the path of the moving heat source based on the length covered during that time: this model will be referred to as D1, see Figure 1. The velocity of the heat source links the time variable to the positional (spatial) variable. Whether or not this Department of Railway Vehicles, Aircraft and Ships, Budapest University of Technology and Economics, Hungary Corresponding author: Ferenc Kolonits, Department of Railway Vehicles, Aircraft and Ships, Faculty of Transportation and Vehicle Engineering, Budapest University of Technology and Economics, Budapest 1521, Hungary. Email: kolonf@iif.hu at RICE UNIV on May 18, 2015pif.sagepub", " This issue is similar to the Pe\u0301clet number used in heat transfer theory in that the flows are defined in the same medium. For the present case, this Pe\u0301clet-like number will be denoted as B. An arbitrary factor may be applied; its definition in this paper follows that of Carslaw and Jaeger.2 Carslaw and Jaeger2 presented analytical solutions that can be directly or indirectly utilized on moving heat sources. There is a two-dimensional model that also includes the heat conduction parallel to the motion, which will be denoted as D2x, see Figure 1. A full three-dimensional solution is also shown, that is denoted as D3, see Figure 1. These formulas are of increasing complexity and are unsuitable for application in analytical studies. Therefore, their relationships should be evaluated on the basis of numerical results. For comparative computations, a simple model has been chosen: a heat source of heat flux intensity q distributed uniformly on a (2w) (2l)-sized rectangular area travelling at a velocity v straight along the border of the half-space, all values being constant. It is known that the temperature distribution depends on the actual form of the heat source", " A systematic application with D1 has already been made by Blok, and also with D2x for some measures. The aim of the present study is to comparatively evaluate the models with sufficiently detailed numerical calculations. D2y will presumably give a better approximation to D3 than do D1 and D2x. Under what circumstances this is in fact the case is the essential question considered in the present study. Models for a moving heat source on a half-space with various conduction properties The surface temperature is considered based on Figure 1. The x\u2013y in-plane coordinates are made dimensionless using the heat source size as , i.e. the source lies at 14 , 4 1. Case D1. This formula (being \u2018Blok\u2019s original\u2019 formulation) is widely known and is obtained by integrating the one-dimensional heat spike formula over time and then transforming the result into a moving coordinate system (see, for example, p. 269 of Carslaw and Jaeger2 and also Vernersson4) TD1 \u00f0 \u00de \u00bc 2qw l ffiffiffiffiffiffiffiffiffi 2B p ffiffiffiffiffiffiffiffiffiffi \u00fe 1 p ffiffiffiffiffiffiffiffiffiffi 1 ph i \u00f01\u00de The square-bracketed term takes effect if > 1; it shows that the heat input has stopped and the temperature rise pertaining to it must be subtracted from the initial expression" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001227_s10846-014-0160-4-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001227_s10846-014-0160-4-Figure1-1.png", "caption": "Fig. 1 Attitude configuration of the unmanned helicopter", "texts": [ " The rest of this paper is arranged as follows. In Section 2, the mathematical model of the unmanned helicopter is given. In Section 3, the nominal controller and robust compensator design procedure is proposed. In Section 4, the properties of designed controller are discussed. In Section 5, simulation and real flight tests are carried out, with experimental results analyzed and evaluated. Finally, in Section 6, conclusions are given. The helicopter configuration adopted in this paper is shown in Fig. 1, and main variable descriptions are listed in Table 1. In general, the dynamics of the helicopter are nonlinear with coupling effects [13]. Linearized models, in which nonlinearities and parameter uncertainties are ignored, have been adopted in much of the available literature, e.g., in [2, 6, 9\u201313]. In this paper, a linear model, in which parameters can be measured precisely, is adopted as the nominal model. However, the uncertain parts, including nonlinearities, parameter uncertainties, external disturbances, and other unmodeled dynamics, are not ignored, all being contained within equivalent disturbance" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002673_j.matpr.2020.07.229-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002673_j.matpr.2020.07.229-Figure5-1.png", "caption": "Fig. 5. Contour plot of 3rd Mode shape (Twisting mode) of cantilever beam.", "texts": [], "surrounding_texts": [ "While conducting the model analysis of the composite cantilever beam, convergence criteria was achieved after several iterations. Using an iterative method, we increase the number of elements by decreasing the size of elements. To determine the suitable mesh size which produces a precise result, the mesh size was decreased until the natural frequencies for a mode shape converged. For the modal analysis of the beam, the number of nodes obtained were 11,028 and the number of elements were 10,983 for an element size of 3 mm. a meshed model of the beam is shown in Fig. 2." ] }, { "image_filename": "designv11_22_0002179_icuas.2016.7502682-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002179_icuas.2016.7502682-Figure1-1.png", "caption": "Fig. 1. Quadrotor model", "texts": [ " Thus, singularities of local parameterization are completely avoided to generate agile maneuvers in a uniform way, (iv) the proposed algorithm is validated with numerical simulations along with real-time experiments with a quadrotor UAV. This paper is organized as follows. A dynamic model is presented and problem is formulated at Section II. Control systems are constructed at Sections III and extended kalman filter developments are presented in IV, which are followed by numerical examples in Section V and an experimental results in Section VI. Consider a quadrotor UAV model illustrated in Figure 1. We choose an inertial reference frame { e1, e2, e3} and a body-fixed frame { b1, b2, b3}. The origin of the body-fixed frame is located at the center of mass of this vehicle. The first and the second axes of the body-fixed frame, b1, b2, lie in the plane defined by the centers of the four rotors. The configuration of this quadrotor UAV is defined by the location of the center of mass and the attitude with respect to the inertial frame. Therefore, the configuration manifold is the special Euclidean group SE(3), which is the semi-direct product of R 3 and the special orthogonal group SO(3) = {R \u2208 R 3\u00d73 |RTR = I, detR = 1}", " Figures 4(b), 4(d), and 4(f) illustrate the angular velocity of the quadrotor during this maneuver. Noisy measurement data and the estimated value from EKF are presented along with the true path to show the performance and effectiveness of the EKF to reduce the error and improve noisy measurements. Position and velocity estimation errors for this example are presented in Figures 4(g) and 4(h). The quadrotor UAV developed at the flight dynamics and control laboratory at the George Washington University is shown at Figure 5(a). We developed an accurate CAD model as shown in Figure 1 to identify several parameters of the quadrotor, such as moment of inertia and center of mass. Furthermore, a precise rotor calibration is performed for each rotor, with a custom-made thrust stand as shown in Figure 5(b) to determine the relation between the command in the motor speed controller and the actual thrust. For various values of motor speed commands, the corresponding thrust is measured, and those data are fitted with a second order polynomial. Angular velocity and attitude are measured from inertial measurement unit (IMU)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001203_s12541-014-0415-9-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001203_s12541-014-0415-9-Figure3-1.png", "caption": "Fig. 3 Gear after test", "texts": [ " For the gear contact fatigue performance, in order to fully understand the tooth contact fatigue life, the traditional group method was used in this test which adopted four groups of test loads and five samples for each test load. It is ensured that the gear does not have the bending fatigue broken teeth during the test. After the test, scanning electron microscopy (SEM) was used to determine the pitting failure shape and size. Two different SEM analyses were carried out as macro and micro SEM analysis. Macro analysis technique was used to determine the relationship of pitting failures, and micro analysis was used to study the fracture tip and nodule-crack relationship. After the test, as shown in Fig. 3, it is verified that the pitting failures occur between the initial point of single tooth contact and pitch line. SEM figures of gear contact fatigue failure on the tooth surface is given in Fig. 4. It is easy to observe that the three phases of gear fatigue failure, namely crack initiation, crack propagation, Crack instable propagation (spalling). According to the observation of test gear tooth surface and the location of fatigue crack, the test fatigue can be divided into two kinds of fatigue spallings which are caused by subsurface fatigue failure and surface fatigue failure" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003627_icedme52809.2021.00056-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003627_icedme52809.2021.00056-Figure1-1.png", "caption": "Figure 1. FFrobot-III robot", "texts": [ " The world's first wall robot prototype adopts negative pressure adsorption. It was developed by Lecturer Xiliang of Osaka Prefecture University in Japan in 1966. The air in the suction cup was sucked by a ducted fan to achieve the purpose of negative pressure adsorption. In 1975, In 1982, wheeled and walking-footed negative pressure wall robots were developed separately[3]. With the development of negative pressure technology, in China, the FESTO Laboratory of Chongqing University developed FFROBOT-III in 2018[4], as shown in figure 1. Compared with the previous FFROBOT-I and FFROBOT-II, it uses more vacuum Sucker adsorption, its moving platform is changed from H-type parallel mechanism to a cross-type mechanism. The horizontal and vertical directions are driven by their respective motors independently, and the control is independent and does not affect each other. There is no synchronization problem of the motors. At the same time, the cross-type mobile platform is improved In the mobile space, while ensuring a slight increase in the step distance, the frame size is greatly reduced, effectively reducing the self-weight caused by the frame" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002798_s00170-020-06152-6-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002798_s00170-020-06152-6-Figure12-1.png", "caption": "Fig. 12 Use of three inspection oriented datum target features (A1,2,3) for datum A establishment", "texts": [], "surrounding_texts": [ "The well-known \u201cmaximum utilization\u201d rule concerning datum hierarchy states that \u201cIf a Datum Feature can and may constrain a degree of freedom, it must.\u201dWhen a DF is used to establish a datum, it constrains, or locks, some degrees of freedom of an ideal feature (e.g., the tolerance zone). As per ISO 5459, the maximum number of degrees of freedom that can be constrained by this integral feature is equal to, or less than, six minus the invariance degree of the nominal integral DF. However, high level of flexibility is again here provided to the designer, concerning the designation of the number of locked or released degrees of freedom in each datum of a datum system. This can be specified by the assignment of complementary indication ([PL], [SL], [PT], ><, [ ] or [Tx,Ty,Tz,Rx,Ry,Rz]) added after the datum identifier symbol in the relevant datum section, as shown in the example given in Fig. 7." ] }, { "image_filename": "designv11_22_0000929_romoco.2015.7219726-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000929_romoco.2015.7219726-Figure1-1.png", "caption": "Figure 1. Kinematic scheme of the mobile manipulator.", "texts": [ " Multiplying the dynamic equation (8) by qN T the vector of Lagrange multipliers is eliminated and finally the new form of dynamic model is obtained: BuqNqqFqNqqMqN TTT , where rrkpr rp p qN qN 10 0 ~ is knn dimensional full rank matrix. Determining controls u from (18) and substituting into conditions (7) system of linear inequalities, underdetermined with respect to gain coefficients is obtained. The solution of this system are parameters I iV , , I iL , , II iL , ensuring fulfillment of constraints (7) at each time instant. Details of this method are presented in [14]. In the numerical example, a mobile manipulator, shown in Fig. 1, consisting a nonholonomic platform of (2,0) class and a 3DOF RPR type holonomic manipulator working in a three-dimensional task space is considered. The mobile manipulator is described by the vector of generalized coordinates: T cc qqqyxq 32121 ,,,,,,, where cc yx , denotes the platform center location and is the platform orientation, 21 , are angles of driving wheels, 1 q , 2 q , 3 q are angles and offset of the manipulator joints. The platform works in BB YX plane of the base coordinate system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003429_j.tws.2021.107585-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003429_j.tws.2021.107585-Figure2-1.png", "caption": "Fig. 2. General geometry of the cylindrical pressure vessel.", "texts": [ " [36], indicating the equivalency between torispherical and ellipsoidal ends is not supported by the finite element method analyses and the membrane theory. In this paper ellipsoidal dished end is assumed to be the one described by the true elliptical shape. Within the framework of this study, the dished ends are defined with a reference to the geometry of their middle surfaces, which is preferable to perform further mathematical description. The general geometry of the cylindrical pressure vessel is presented in Fig. 2. The relative depth of a dished end is defined as \ud835\udefd = \u210e \ud835\udc450 . (5) The parameter \ud835\udefd is similar to the inverse of \ud835\udc3e (Eq. (3)). In the paper the case of \ud835\udefd = 0.5 is analysed, as it is the most common value for the pressure vessels. The principal radii of curvature \ud835\udc451, \ud835\udc452 are related with parallel radius \ud835\udc5f in the following manner: 1 = d\ud835\udc5f d\ud835\udf11 1 cos\ud835\udf11 , \ud835\udc452 = \ud835\udc5f 1 sin\ud835\udf11 . (6) To perform the necessary calculations using the Ritz method, four separate shells have to be considered i.e. toroidal, spherical, ellipsoidal and cylindrical", " (8) a p s r \ud835\udf11 w \ud835\udf11 t \ud835\udf11 w t \ud835\udf09 3 3 t t s \ud835\udf00 w z \ud835\udc51 i p \ud835\udf0e S \ud835\udf0e In the case of ellipsoidal dished end the principal radii of curvature depend on the angle \ud835\udf11: \ud835\udc451\ud835\udc52 = \ud835\udc450 \ud835\udefd2 sin3 \ud835\udf11 (1 + \ud835\udefd2 cot2 \ud835\udf11)\u2212 3 2 , \ud835\udc452\ud835\udc52 = \ud835\udc450 1 sin\ud835\udf11 (1 + \ud835\udefd2 cot2 \ud835\udf11)\u2212 1 2 . (9) Finally, the geometry of the cylindrical shell is described as follows: \ud835\udc451\ud835\udc50 \u2192 \u221e, \ud835\udc452\ud835\udc50 = \ud835\udc450. (10) Each of the shells is described by individual range of a parameter long which its geometry is defined. For the doubly curved shells, the arameter is meridional angle \ud835\udf11 shown in Fig. 2. For the torispherical, pherical and ellipsoidal shell the parameter changes in the following anges: \ud835\udc611 \u2264 \ud835\udf11 \u2264 \ud835\udf11\ud835\udc612, \ud835\udf11\ud835\udc601 \u2264 \ud835\udf11 \u2264 \ud835\udf11\ud835\udc602, \ud835\udf11\ud835\udc521 \u2264 \ud835\udf11 \u2264 \ud835\udf11\ud835\udc522, (11) here: \ud835\udc601 = \ud835\udf11\ud835\udc521 = 0, \ud835\udf11\ud835\udc602 = \ud835\udf11\ud835\udc612 = \ud835\udf11\ud835\udc522 = \ud835\udf0b 2 . (12) The parameter \ud835\udf11\ud835\udc611 = \ud835\udf11\ud835\udc602 has to be determined. Following the descripion in the standard [4] it is possible to estimate that \ud835\udc611 = \ud835\udf11\ud835\udc602 \u2248 3 20 \ud835\udf0b. (13) The geometry of the cylindrical shell is defined with a use of coordinate \ud835\udf09, assuming: \ud835\udf09 = \ud835\udc65 \ud835\udc3f , 0 \u2264 \ud835\udc65 \u2264 \ud835\udc3f. (14) here \ud835\udc3f \u2014 length of the cylindrical shell" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000353_iemdc.2019.8785381-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000353_iemdc.2019.8785381-Figure3-1.png", "caption": "Fig. 3. Example of a 2D slice, used for the skewing analysis. In this case, the slice corresponds to \u03b1sl = \u22122/5\u03b1sk", "texts": [ " A certain number of 2D slices, of the continuous rotor skewing, is considered. If ntot is the number of slices, the angle of each slice, with respect the RFO dq reference frame, set in the middle section of the machine, is: \u03b1sl = \u03b1sk 2n\u2212 ntot \u2212 1 2ntot (7) where n = 1, . . . , ntot is the considered slice and \u03b1sk is the skewing angle. In the following example, ntot = 5 slices have been considered. Using static simulations, the condition to have the same current in corresponding bars, in the several slices, is automatically imposed. In Fig. 3 a resolved 2D slice is reported. In this case the rotor and its electric load are rotated forward. During the skewing analysis the dq reference frame is fixed, as well as the angle between it and the rotor \u03b1\u03b2, despite several rotor positions are analyzed. The rotor current computation depends upon the inductances variation due to skewing. Thus, during the iterative procedure, for each value of the rotor current, the multislice skewing analysis is included, in order correctly evaluate the rotor qaxis flux linkage" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003547_j.jmapro.2021.04.060-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003547_j.jmapro.2021.04.060-Figure2-1.png", "caption": "Fig. 2. The finite element model for distortion prediction.", "texts": [ " Thus, the initial condition is applied as: T(x, y, z, 0) = Tb (7) The air natural convection and radiation to the ambient environment on the top surface of the powder bed during the AM process can be expressed by [21]: \u2212 k \u2202T \u2202\u03bc = \u03b4\u03c9 ( T4 \u2212 T4 \u221e ) + c(T \u2212 T\u221e) (8) where \u03bc is the unit normal vector, \u03b4 is the Stefan-Boltzmann constant, \u03c9 is the emissivity, T\u221e is the ambient temperature, and c is the heat convection coefficient. Regarding the mechanical aspect, the substrate is fixed. The interaction of the substrate and as-printed specimen is realized by the entire tie-constraint, which is the identical way of contacts between adjacent layers. This mimics the boundary conditions of what happens in reality. The thermal-recrystallization-mechanical model is implemented through the Abaqus\u00ae. Various scanning paths and the heat source model are achieved by the corresponding subroutine program [22]. As exhibited in Fig. 2, a monolayer rectangle sample (15 mm \u00d7 3 mm \u00d7 0.1 mm) with a rectangular substrate is adopted to scrutinize various scanning strategies. The average mesh size of the deposition layer is 0.3 mm \u00d7 0.3 mm \u00d7 0.1 mm, and the total number of elements is 1500. As the mesh size is smaller than the diameter of the laser spot 0.6 mm, the computed results are reasonable [12]. The relatively coarse mesh is employed in regions away from the zone of interest. The analysis steps of all the simulations are in automatic form" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure1-1.png", "caption": "Fig. 1. Configuration of Astraios.", "texts": [ " Furthermore, it can be applied to pitch and yaw bearings of various sizes and to the performance test of pitch and yaw drives. The applicability of the test rig to 2.0- 3.0 MW class wind turbines was shown through structural analysis and preliminary friction torque tests. Only a few test rigs have been developed for testing large *Corresponding author. Tel.: +82 42 868 7994, Fax.: +82 42 868 7477 E-mail address: yjpark77@kimm.re.kr \u2020 Recommended by Editor Sung-Lim Ko \u00a9 KSME & Springer 2014 wind turbine bearings. Among them, Astraios from Shaeffler Technologies is the best known one (Fig. 1) [7]. This rig comprises eight cylinders\u2014four radial and four axial\u2014that can be used to represent 6 degree of freedom (DOF) loads. Various types of bearings, including general slewing bearings and main bearings, can be tested at rotation speeds of 4-20 RPM through the drivetrain connection. The testable maximum outer diameter is 3.5 m. However, this rig is more suitable for testing main bearings than for testing pitch and yaw bearings; this is because the drive type used is a drivetrain that is suitable for main bearings" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002665_s11071-020-05846-6-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002665_s11071-020-05846-6-Figure1-1.png", "caption": "Fig. 1 Main input\u2013output characteristics of the tire cornering mechanical behaviors", "texts": [ " In addition to the tire cornering characteristics, the tire mechanics models also include the tire longitudinal mechanical model, the tire vertical vibration model and the tire model in combined working conditions [39]. Different tire models are used to analyze the corresponding vehicle dynamics characteristics and establish the vehicle dynamics models for subsequent control design. The main input\u2013output characteristics of the tire cornering mechanical behaviors, which are studied in this work, are shown in Fig. 1. It can be seen that the lateral forces generated by the tire in vehicle turning process are mainly influenced by the tire slip angle, the tire vertical load and the road adhesion coefficient, among which the first two factors are changed dynamically according to the vehicle actual driving process, while the third factor is only related to the road. For conventional vehicle lateral dynamics control design, which is assumed to be researched when the vehicle is turning with small angle, thus the relationship between the tire lateral force and its influencing factors can be regarded as linear [40, 41]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003240_j.engfailanal.2020.105208-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003240_j.engfailanal.2020.105208-Figure16-1.png", "caption": "Fig. 16. Pre-deformation angle.", "texts": [ " In this study, the pre-deformation angle, initial tightness, and roughness are chosen as the design variables as they have a negligible impact on the vibration characteristics and performance of the blade. The limits of the three design variables are listed in Table 2. The pre-deformation angle in the cold shroud airfoil, assuming that the working condition deformations along the axial and tangential directions of the contact interface are different, is represented by \u03b8. This can be eliminated by setting a pre-deformation angle during the assembly stage. The pre-deformation angle is shown in Fig. 16, with a limit of Y.-X. Liu et al. Engineering Failure Analysis 122 (2021) 105208 \u00b15\u25e6. The initial tightness of the contact surface is represented by t, with a value of 0.05\u20130.25 mm. The friction factor, which indicates the effect of roughness, is represented by u, with a value of 0.15\u20130.35. The target of the design optimization process is the contact stress. Y.-X. Liu et al. Engineering Failure Analysis 122 (2021) 105208 About 70 sampling points were calculated within the design space, and the parametric sensitivity was obtained, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002491_1077546320932030-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002491_1077546320932030-Figure2-1.png", "caption": "Figure 2. Test rig. (a) Schematic of load devices of machinery system and (b) actual experimental system.", "texts": [ " To verify the effectiveness of CIKECA algorithm, two experimental studies on fault diagnosis of rolling bearings were carried out, and details of the setup are described as follows. Where Experiment 1 is mainly aimed at different types of rolling bearing state recognition. And Experiment 2 mainly studies the influence of different noises (i.e. the distance from the fault vibration source) on the state identification rate of rolling bearing. The vibration measuring system is mainly composed of hardware and software part, and the schematic of load devices of machinery system and the actual experimental system are shown in Figure 2(a) and (b), respectively. Deep groove ball bearings (6328-2RZ) are applied as the test bearing. The geometric parameters of the bearings are listed as follows: inner ring diameter (Di) = 28 mm, outer ring diameter (Do) = 28 mm, ball diameter (Db) = 11.509 mm, pitch diameter (Pd) = 48.5 mm, number of balls (Z) = 8, and the contact angle (\u03b1) = 0.274 rad. The inner ring rotates uniformly with the shaft driven by the driving device, whereas the outer ring maintains a static state under the central axial load" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001221_esda2014-20001-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001221_esda2014-20001-Figure1-1.png", "caption": "FIGURE 1. IN-SITU MOUNTING OF THE LASER CLADDING TECHNOLOGY", "texts": [ " Direct, laser-based metal deposition is a novel concept for the fabrication and repair of components as well as geometrical surface modifications. This analysis clearly demonstrates that laser cladding can be applied to in-situ marine diesel engine crankshaft repairs. The conceptual model of the laser cladding machine for in-situ crankshaft repairs has therefore been developed. The major goal of this research project is to design a fullscale functional prototype of the in-situ laser cladding apparatus (hereafter \u2013 device) for renovation and repair of crankshaft journal surfaces (see Fig. 1). The initial idea of this application is outlined in the Patent of the Republic of Latvia no. B24B5/42 - Device and method for the in-situ repair and renovation of crankshaft journal surfaces by means of laser build-up [14]. The principal design of the in-situ laser cladding device which is proposed in this article is subject to the International Patent application PCT/LV2013/000006 of 18.07.2013 \u2013 Apparatus and method for in-situ repair and renovation of crankshaft journal surfaces by means of laser cladding [15]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001886_we.2008-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001886_we.2008-Figure8-1.png", "caption": "Figure 8. The four rotor positions that represent each set of DLC 2.3 simulations. The dotted figure represents a reference position, and it is plotted for illustration purposes only. The gray blade is to ease the visualization of the rotation.", "texts": [ " An example of the EOG is shown in Figure 7, along with the time of occurrence of the fault. Following the guidelines of the IEC standard, the event is simulated at 11.4 m/s (rated wind speed), 9.4 m/s, 13.4 m/s and 25 m/s (cut-out). In addition, the starting azimuth of the wind turbine rotor is changed for each set of simulations. That is, the initial orientation of the blades is changed by 60\u00b0. This is carried out so at the gust occurrence, several states of the rotor position are considered adding randomness to the process. The rotor positions considered here are presented in Figure 8, and the numeric data specifying the detail positions are presented in Table IV. In summary, a total of 48 simulations will be used to study the extreme loads in the gearbox bearings. Wind Energ. (2016) \u00a9 2016 John Wiley & Sons, Ltd. DOI: 10.1002/we Table IV. Angles (in Deg.) of the blades in the different rotor positions, where 2 is the gray shadowed blade in Figure 8. The phase angle between each position is 30\u00b0. Case 1 2 3 Position 1 180 60 60 Position 2 210 90 30 Position 3 240 120 360 Position 4 270 150 30 Figure 9. The system\u2019s block diagram used for the emergency brake load case. Certain situations, such as excessive tower-top vibrations or when human safety is at jeopardy, require the stop of the wind turbine abruptly using an emergency stop procedure. In this case, it is a decision taken by the turbine operator or a pre-defined safety feature of the wind turbine controller" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002762_j.euromechsol.2020.104125-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002762_j.euromechsol.2020.104125-Figure4-1.png", "caption": "Fig. 4. (a) Configuration during rolling element interaction with inner race defect, (b) sketch depicting various angles.", "texts": [ " (12)(Chen and Kurfess, 2019) [ Lom cos\u03c8 spall \u2212 Los cos\u03c8d ]2 + [ Lom sin\u03c8 spall \u2212 Los(sin\u03c8d+sin\u03c8di) ]2 =(0.5Db) 2 (12) where the angle \u03c8dican be obtained from Eq. (13) wherein tdi is found out from the signal \u03c8di =\u03c9cagetdi (13) Using the angle \u03c8spallobtained from Eq. (12) the length of the spall is computed using Eq. (14) Lspall = \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 1 \u2212 cos ( \u03c8 spall ) 2 \u221a ( Dp +Db \u2212 cl \u2212 2\u03b4 ) (14) The expression for spall length for a spall on the inner race can be derived in a similar fashion. The sketch of the interaction of the rolling element with inner race spall is shown in Fig. 4. Similar to the outer race defect, the time separation from start to complete destressing in case of inner race defect is given in Eq. (15) Tsd = \u03b1b + \u03b1d \u03c9cage (15) where, the angle \u03b1b can be expressed by Eq. (16) using simple trigonometry as shown in Fig. 4(b) \u03b1b \u2248 tan(\u03b1b)= bi( 0.5Dp \u2212 0.5Db ) \u2212 (\u03b4o + 0.5cl) (16) where, biis semi-width of contact area in tangential direction (Figs. 2(d) and Fig. 4(b) which can be expressed similar to Eq. (3) as (Harris, 1996) bi = b* i [ 3F\u03c6 2 \u2211 \u03c1i ( 1 \u2212 \u03be2 I EI + 1 \u2212 \u03be2 II EII )]1 / 3 (17) where, b* i is dimensionless parameter dependant on curvature of bodies in contact, \u2211 \u03c1i is sum of curvature of inner race-rolling element contact. Now, the distance between the bearing center \u2018O\u2019 and rolling element center \u2018S\u2019 is expressed as LOS = \u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305\u0305 ( 0.5Dp \u2212 0.5Db \u2212 \u03b4o )2 + ( \u03c9cageDpTed 2 )2 \u221a (18) Using Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure6-1.png", "caption": "Figure 6. Creation of the model on ANSYS WB11", "texts": [ "The contact stiffness for an element of area\ud835\udc34is calculated using the following formula (Mohr, 1980): { }( ){ }Tkn i iF f e f dA= \u222b (6) Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. The default value of the contact stiffness factor (FKN) is 1, and it is appropriate for bulk deformation. If bending deformation dominates the solution, a smaller value of FKN = 0.1 is recommended. In this work, a penalty method has been selected to solve the problem of contact. To begin the study, we created a structure on ANSYS Wb represents the brake with the pad (Fig.6). Then we made the mesh and were defined boundary conditions to put on ANSYS Multiphysics and to initialize the calculation. The principle of the model in the literature generally taking into account at every moment the evolution of the contact disc-pad. This distribution of the contact can be calculated and applied heat flux created by friction.In this study, initial mechanical calculation aims at determining the value of the contact pressure (presumably constant) between the disc and the pad. It is supposed that 60% of the braking forces are supported by the front brakes (both rotors), that is to say 30% for a single disc (Mackin et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001569_s40846-015-0064-1-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001569_s40846-015-0064-1-Figure2-1.png", "caption": "Fig. 2 Exterior of video-assisted tactile sensor", "texts": [ " If the difference is more than 5 %, the bisection method is used to determine a new Young\u2019s modulus. For example, if the simulated contact area is larger than the experimental one for a fixed load, it means that the simulated object is softer and thus the Young\u2019s modulus should be set higher. A simulation is then conducted using the new value. This process is repeated until the difference is sufficiently small. The developed device consists of a camera, a dual force sensor, and a sensor head. The corresponding experimental setup is shown in Fig. 2. The force sensor is a built-in component of the employed universal tribometer (UMT3MT, CETR, USA). The tribometer allows continuous acquisition of the normal force, tangential force, indentation distance, and the sliding distance. Figure 3 shows the dimensions of the sensor head. The sensor head is a transparent, elastic, and hemispherical membrane made of polydimethylsiloxane (PDMS). Its elastic modulus was found to be 2100 kPa (mixing ratio: 10:1; curing time: 12 h; temperature: 90 C) using a M", " The porcine liver was assumed to be nearly incompressible (m = 0.499). Quadrangles (QUAD (4)) are set for all elements. The meshes in the contact region are refined to improve simulation accuracy. The object is 40 mm in diameter and 10 mm in depth. To determine the efficiency of the proposed approach, a PDMS standard was first tested. The PDMS standard had a Young\u2019s modulus of 2.1 MPa and a Poisson\u2019s ratio of 0.499 [28]. The cylindrical PDMS block was 60 mm in diameter and 10 mm in height (see Fig. 2). The sample was colored red to increase the contrast in the contact image. The load in the experiments was gradually increased from 0.5 to 3.0 N in steps of 0.5 N. The indentation speed was 0.3 mm/s. At each step, the load was maintained for 15 s and then the contact image was acquired. The ambient temperature was 23 C. The contact area was calculated using ImageJ. Experiments were repeated three times at three locations on the PDMS standard. The contact area corresponding to a given load was averaged" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002750_1.g005244-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002750_1.g005244-Figure1-1.png", "caption": "Fig. 1 Two-dimensional impact-angle-constrained engagement geometry [2].", "texts": [ " This remark is intended to avoid possible misleading or confusion that might arise in the following analyses of various guidance laws, as there is no closed-form expression of the SDREguidance command for each design considered in this section, namely, Eq. (11) in [2] andEq. (51) in [9]. Likewise, the same consideration applies to the guidance design using the SDDRE scheme, Eq. (69a) in [1]. This research is benchmarked by [2] with respect to a pursuer\u2013 evader differential game, and the engagement geometry is illustrated in Fig. 1. The dynamic equations are _x1 \u22122 Vr\u2215r x1 \u2212 cos \u03b1 \u2212 \u03b8 \u2215r uI cos \u03b2 \u03b8 \u2215r uT; _x2 1\u2215VI uI 1\u2215VT uT 9 where x1 (resp., x2) is the line-of-sight (LOS) rate (resp., error between the actual and desired impact angles); Vr (resp., r > 0) denotes the relative velocity (resp., distance); \u03b1 (resp., \u03b2) flight-path angle of the pursuer (resp., evader); \u03b8 LOS angle; uI (resp., uT) control input/acceleration of the pursuer (resp., evader); and VI (resp., VT) pursuer (resp., evader) velocity. According to the design [2], the primary guidance objectives (zero miss and desired impact angle) are so formulated as the stabilization problem of system (9)", " Recently, [1] also contributes to the impact angle guidance problem, novelly in the finite-time framework that is extended from the SDRE scheme (its infinite-time counterpart), namely, the SDDRE scheme, which is used to stabilize the nominal system in the guidance design [1], whose (error) dynamics are _ea 2 664 _x1 \u2212 _xd1 \u2212 _xd2 \u2212 2_rx2\u2215r \u2212\u03b7z 3 775 2 64 0 \u2212 cos \u03b1 \u2212 \u03b8 \u2215r 0 3 75u (15) where ea x \u2212 xd; z \u2208 R3; x1 (resp., x2) is theLOSangle (resp., rate); xd xd1 ; xd2 T \u2208 R2 is the desired trajectory of x; both the dummy variable z \u2208 R and 0 < \u03b7 \u226a 1 are introduced to account for any state-dependent bias term such that the classical SDC construction is applicable [3];udenotes the control input/missile acceleration; and the engagement geometry can also be inferred from Fig. 1. Following the guidance design [1], the stabilization of the output ea 1\u22362 in the control system (15) satisfies the primary guidance objectives. Specifically, x1 \u2192 xd1 to achieve the desired impact angle, whereas x2 \u2192 xd2 to null the LOS rate (_\u03b8 \u2192 0) [1]. However, the theoretical and practical results for this guidance design [1] (including the stability analysis and effectiveness validation) rely on the fundamental assumption of SDDRE\u2019s overall applicability (Assumption 2 in [1]). Therefore, this subsection fulfills such a shortage/theoretical basis, in alignment with the main focus of this Note", " Specifically, there is no need to introduce any excessive dummy variable, x3 in this application [1], which thereby resolves the issue of state-dependent bias term and conforms to the classical SDC construction [3]. In otherwords, the philosophy ofAlgorithm 1 in [13] has already integrated such an issue into the design by exploiting the SDC flexibility, which thus avoids increasing the system dimension or computational complexity via the classical way [3]. 2) So far, the two designs [1,2] in Sec. III.A (SDRE) and Sec. III.B (SDDRE), respectively, are with respect to the 2D engagement geometry (Fig. 1). As extensively validated via MATLAB, the proposed new analytic results essentially and consistently support the practical achievements in the state-of-the-art of SDRE- and SDDREbased impact angle guidance literature [1,2]; namely, the applicability guarantees for the effectiveness validations therein. With such confidence, we thus generalize the analysis scope from 2D to 3D, as presented in Sec. III.C. Moving forward from the preliminary 2D engagement scenario [1,2], this Sec. III.C considers the more realistic 3D impact-angleconstrained guidance law design, and Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001353_1.2369460-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001353_1.2369460-Figure6-1.png", "caption": "Fig. 6. Drawing of details of Mk 40 spur-gear planetary transmission.", "texts": [ " T h e sensing ins t rument used throughout the investigation was a Western Electric 640AA microphone, operat ing on the condenserd iaphragm principle, and designed for frequencies f rom 1 to 30 000 cycles. M k 40 T R A N S M I S S I O N T h e Mk 40 gear train is a singlestage planetary system with straight spur teeth. A central pinion running at turbine speed drives three planets on fixed centers, which in turn deliver the torque to a single internal- toothed ring gear connected to the ou tpu t shaft. Figure 6 shows the specifications for the gears. T h e ratio of the machine is 3.09756 to l. After having been run, the teeth measured to an accuracy within 0.0002 in. in the shape of the involute curve, the tooth spacing, and the concentricity. An occasional error of 0.00025 in. was found. These are gears of great accuracy. T h e transmission was isolationmounted and supplied with cooling oil in the form of a fine internal spray. Its rotat ing parts and shafts were balanced precisely and the machine was lined up to the drive within 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002206_s00170-020-04924-8-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002206_s00170-020-04924-8-Figure6-1.png", "caption": "Fig. 6 a Extraction of middle square contour points at a distance of 0.5 mm from the top. b Points extracted", "texts": [ " b Part\u2019s engineering drawing parameters of 0.01 mm for size and 1 mm for sag value. The part was then manufactured under Replicator 2X FDM printer using polylactic acid (PLA) material. The manufacturing parameters set are shown in Table. 1. The part was then scanned by Atos Core 3D scanner, which offers a measurement precision of 20 \u03bcm, and treated using GOM inspect software. The points constituting the contour of each geometric feature were extracted at a height of 0.5 mm from the top of features (Fig. 6) to avoid the edges that can be rounded which may result in inaccurate measurements, thereafter. The extracted points were then represented by their (x, y, z) coordinates in CCS. As the interest was only in in-plane deviation, only x and y coordinates were retained. These coordinates must be classified from the smallest to the largest (according to x or y optionally) to allow a good conversion to polar coordinates (\u03b8i, rm(\u03b8i)). A MATLAB code to sink our models to the measured data obtained from the manufactured part, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001029_ac60180a049-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001029_ac60180a049-Figure1-1.png", "caption": "Figure 1. Borosilicate flasks for oxygen combustion procedures", "texts": [ " N ACCORDANCE with the practice I followed in previous reports of the Committee on Microchemical Apparatus (1, 9, 4, these specifications are for pieces of apparatus that are either the most widely used in their respective fields of application or are an improvement over such apparatus according to tests made by members of this committee and cooperating chemists. In this report, specifications are recommended for the conventional apparatus used in connection with oxygen flask combustion procedures (7, 8). Types of apparatus employing electric ignition (2, 6, 6) are not included be- cause of lack of experience with these forms. Figure 1 shows two sizes of borosilicate flasks-the 300- and 5Wml. sizes-and a glass stopper with platinum sample holder which is used with a flask of either size. Although oxygen flask combustion VOL. 33, NO. 12, NOVEMBER 1961 1789 has been proved to be generally Mfe, precautions such as the use of gloves, shields, etc., should be taken. (3) Committee on Microchemical Apparatus, Division of Analytical Chemistry, (5) Juvet, R. S., Chiu, J., I&id., 32, 130 (1960). 31, 1932 (1959). 74; 1955, 123; 1956,869" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002408_s11012-020-01162-w-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002408_s11012-020-01162-w-Figure3-1.png", "caption": "Fig. 3 Kinematic details of the present two-section wing", "texts": [ " In the field of the kinematic science, position, velocity, acceleration, and all higher derivatives of the space variables (relative to time or any other variables) are investigated, and accordingly, the kinematics of the connected mechanical robotic manipulators contains all of the geometric and the time dependent features of the motion. In order to extract these kinematic equations, the location and orientation of the connected links are investigated. In order to simplify the study of their movement, the frames are connected to the link joints according to the Fig. 3. The lengths of the first and second sections of the wing are respectively L1 and L2, and the flapping angles of each section around their respective axes are c1 and c2 respectively. The wing first section has the constant chord of c1,i=1,\u2026,5 = 0.2, and the second section has the variable chord of c2,i=2,4,\u2026,10 = c2,i=1,3,\u2026,9 = (0.2,0.25,0.2,0.15,0.1). According to Fig. 3: r1 \u00bc s1 cos c1 s1 sin c1 \u00f038\u00de v1 \u00bc s1 _c1 sin c1 s1 _c1 cos c1 \u00f039\u00de r2 \u00bc L1 cos c1 \u00fe s2 cos\u00f0c1 \u00fe c2\u00de L1 sin c1 \u00fe s2 sin\u00f0c1 \u00fe c2\u00de \u00f040\u00de v2 \u00bc L1 _c1 sin c1 s2\u00f0 _c1 \u00fe _c2\u00de sin\u00f0c1 \u00fe c2\u00de L1 _c1 cos c1 \u00fe s2\u00f0 _c1 \u00fe _c2\u00de cos\u00f0c1 \u00fe c2\u00de \u00f041\u00de In which, r1= [x1,y1] T and r2= [x2,y2] T represents the position vector of each proposed point of the first and the second sections of the wing relative to the origin of the coordinates attached to their joint. v1= [vx,1,vy,1] T and v2= [vx,2,vy,2] T are the corresponding velocity vectors at those points, and s1 and s2 are measured along the span length relative to the origin of the first and the second sections, respectively. In the present model, the effect of aerodynamic forces on the wing, wing torsion, and body angle of attack are considered. The aerodynamic forces are as the distributed forces applied to the wing. Therefore, in addition to the concentrated force applied from the power transmission system, the distributed force that actually represents the aerodynamic forces on the wing, must also be taken into account. According to Fig. 3, Newton\u2019s second law for the bond graph modeling of the distributed aerodynamic forces on the wing and the concentrated forces applied by the mechanism to each sections of the wing can be present as follows: X F \u00bc ma \u00bc X2 i\u00bc1 ZLi 0 fi xi; t\u00f0 \u00dedri \u00fe Xn k\u00bc1 Fi;k xi;k; t 0 @ 1 A k \u00bc 1; 2; . . .;1 \u00f042\u00de where fi(xi,t) presents the distributed force of the wing ith section, and Fi,k(xi,k,t) represents the concentrated force which is applied from the mechanism to the wing ith section at point xi,k, andm and a are respectively the wing mass and acceleration" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002534_012146-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002534_012146-Figure8-1.png", "caption": "Figure 8. Distribution of the magnetic field (axial B) of cross-section of axial yoke of bearing shield", "texts": [ " To visualize the flow paths of the main magnetic flux: consisting of the sum of the fluxes of permanent magnets and the axial magnetic flux of the superconducting field coil, the simulation results are shown in figures 5 through 13. The figures 5-13 shows that the value of magnetic induction in ferromagnetic parts (tooth, yoke) does not exceed the permissible values. The figures 5 through 8 show cross sections of an electric machine in the plane of the first packet (Fig. 5), in the plane of the axial interpacket yoke (Fig. 6), the second packet (Fig. 7), and also between the second packet and the axial excitation coil (Fig. 8). 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10.1088/1742-6596/1559/1/012146 An analysis of the magnitude of the magnetic induction shows that the iron is not in saturation mode - the magnitude of the induction does not exceed 2.1 T. The direction of magnetization of the poles of 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003530_s11740-021-01050-6-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003530_s11740-021-01050-6-Figure6-1.png", "caption": "Fig. 6 Tibia fixation plate used as a case study part", "texts": [ " The inert gas is filtered and reused when the L-PBF machine is processing material, in the meanwhile, the machine uses a constant flow during each fabrication cycle. In this report, we analyze the detailed energy consumption calculations in the production of a tibia implant using an L-PBF machine model Renishaw AM400. The relevant machine specifications are listed in Table\u00a06. (16)%loss = (Pin \u2212 Pout) Ptank 1 3 This case study involves the use of a tibia fixation plate or implant in order to present a well-defined product related with AM real applications (Fig.\u00a06). The objective of the study is to analyze the energy consumption from the production of the part including the support structures. The part has the following dimensions: 32.71 \u00d7 21.23 \u00d7 190\u00a0mm. Other specifications are included in Table\u00a07. In Table\u00a08, the process parameters and material specifications are listed. They have been selected from different screening experiments and taking into account the literature review provided in Table\u00a03. The laser scanning strategy selected for the fabrication of this product was quad island (meander) in which the laser is divided into small squares and fills the area in a different order" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003387_icccr49711.2021.9349369-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003387_icccr49711.2021.9349369-Figure4-1.png", "caption": "Figure 4. Parameters used to describe the initial state.", "texts": [ " In this paper, we aim to design a controller that performs better than the conventional controller in the unmodelable section. To illustrate the ability of two controllers, we choose a balance recovery task, aiming to stabilize the unstable system from a tilt angle (with zero angular velocity) and keep balance for a period of time. Now we will give the parameters to describe the initial state of the robot. For the ballbot standing on the ball with all three omniwheels contacting with ball, the body can be regarded revolving around the center of the ball. As shown in Fig. 4, the orientation of the ball can be described: 1) rotates around the z-axis of the ball 2) rotates around the y- axis of the ball. Generally speaking, the body has 2 DOFs: the inclination angle and the direction to which the robot 1) Simplified mathematical model: A nonlinear 3D model of the system dynamics of ballbot has been analytically derived in [2], where the wheel dynamics are neglected. In this model, the three-dimensional system is approximated in three 3 Authorized licensed use limited to: California State University Fresno" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001985_1.4034768-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001985_1.4034768-Figure16-1.png", "caption": "Fig. 16 Whirling motions of both defective rotors", "texts": [ " As a result, the interaction forces between the defective disk and rods cause the disk\u2019s axis-center to have displacement rb, which is shown in Fig. 14. The position error causes e\u00bc 0.59 lm and it remains unchanged; rb equals 0.28 lm after pretightening and it reaches 1.48 lm at 7500 rpm. Figure 15 shows that the rod-fastened rotor with circumferential position error of rod-holes also has three parameter regions A, B, and C. The rotor bending reduces the system stability. This property still exists that the vibration of rod-fastened rotor decreases first and then increases. Figure 16 shows that the whirling orbits of rod-fastened rotor split further than that of integral rotor. Thus, rod-fastened rotor has larger amplitude at harmonic frequency (131.5 Hz) and smaller amplitude at Hopf T frequency (75.6 Hz), which is shown in Fig. 17. 5.2 Concentricity Error of Disks. The third machining error is the concentricity error of coaxial-locating interfaces, which is shown in Fig. 18. Due to locating inaccuracy, the middle convexity of disk b deviates from the ideal center O to the practical center O0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002489_1089-313x.24.2.59-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002489_1089-313x.24.2.59-Figure1-1.png", "caption": "Figure 1 A, Measurement for lower-limb length in fully extended position from iliac crest to toes with fully plantar flexed ankle; B, starting height in 90\u00b0 squat from iliac crest to ground vertical distance; C, jumping task for obtaining mechanical variables for the F-V profile.", "texts": [ " Each dancer had a minimum of 6 years experience (mean \u00b1 SD = 9 \u00b1 2.6 years), and all were exclusively involved in classical ballet training and performance for the last 4 years. The number of hours per week of choreographic rehearsal and supplemental training are detailed in Table 1. Weight (kg), height (cm), and measures for lower-limb length (cm) in fully extended position from iliac crest to toes with plantar flexed ankle, and starting height in 90\u00b0 squat from iliac crest to ground vertical distance (Fig. 1), were measured prior to any physical testing on the morning of the study. Body weight was measured using a Tanita SC-330 (Tanita Corp., Japan), height was estimated with an aluminum stadiometer (model 713, Seca, Hamburg, Germany), and the lower limb length and starting height at 90\u00b0 were measured using a tape measure.15,16 Two days prior to testing, all participants undertook a familiarization session consisting of a standardized warm up of 10 minutes of jogging, dynamic stretching, and 10 repetitions of countermovement jumps (CMJ) with variously loaded Olympic barbells on their shoulders", " The required measurements for determining the optimal F-V profile during jumping performance are the athlete\u2019s body mass, jump height, and push-off distance (Hpo), as measured by the difference between the lowerlimb length in fully extended position and the starting height at 90\u00b0.14,15,20 At the beginning of the testing session, the dancers performed the standardized warm up with which they had been familiarized the day before. Then, to calculate the actual F-V profile, each dancer performed three vertical maximal CMJs with barbell loads corresponding to 0%, 10%, 20%, 30%, 40%, 50%, and 70% of their own body weight in randomized order with 2 minutes of recovery time between trials and 4 minutes between load conditions to avoid fatigue effect15,29 (Fig. 1). The highest score of the three attempts for each load was selected for the F-V profile analysis and the 70% body weight load was included to determine if the participants were able to jump approximately 10 cm (8.48 \u00b1 1.02 cm), as has been recommended in the literature.16 Jump height and the F-V profile were measured using a smartphone app (My Jump 2 available on the Apple App Store) on an iPad device that featured a camera frame rate of 240 fps.32 My Jump 2 provided information regarding the magnitude and direction of the F-V imbalance for each dancer (F-VIMB), theoretical maximal force (F0), theoretical maximal velocity (V0), and theoretical maximal power (Pmax), according to Samozino\u2019s method" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001240_icuas.2014.6842360-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001240_icuas.2014.6842360-Figure3-1.png", "caption": "Fig. 3. Pure rolling motion", "texts": [ " Fx = q\u0304SCx0 (13) Fy = q\u0304SCy (14) N = q\u0304SCn (15) where Cx0, Cy y Cn are the aerodynamic coefficients involved for the lateral dynamics. These coefficients are obtained by considering small angles and with low airplane speed. The following equations describe the dynamics for the roll motion: \u03c6\u0307 = p (16) p\u0307 = L\u0304 Ixx (17) V\u0307y = Fy m + pVx (18) V\u0307x = Fx m \u2212 pVy (19) where p denotes the roll rate, L\u0304 is the rolling moment, Ixx represents the inertia for the x-axis and \u03c6 describes the roll angle. The aerodynamic effects on the airplane are obtained as they have been obtained in the yaw motion. In the Figure 3, it is observed that \u03b4a represents the deviation of the ailerons. In the case of the roll moment, this corresponds the expression L\u0304 = q\u0304SbCL, where b is the wing span of the airplane and CL represents the aerodynamic coefficient of the roll moment [5]. In this section, we describe the linear controllers that have been designed in order to control the fixed-wing MAV. In order to design the altitude control law, we consider the equations defining the longitudinal dynamics, except the equation (1) which defines the linear longitudinal velocity, because it is considered to be constant; the equation (2) is not used, because it represents the flight-path angular rate" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003485_j.jmatprotec.2021.117165-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003485_j.jmatprotec.2021.117165-Figure1-1.png", "caption": "Fig. 1. Conceptual design of the scanning DED nozzle for high throughput. (A) Schematic diagram of AM process, where a square flat-top beam is applied for a high deposition rate. (B\u2013C) Assumption of waviness formation of metal bead, according to the shape of the outlet channel. As the outlet channel increases, the waviness of the surface of the metal bead decreases, making it easy to overlay layer by layer. (D\u2013E) Deposition efficiency according to the distance between the outlet and the surface of the base material. The shorter the distance, the greater the amount of metal powder concentrated within the laser beam which increases the deposition efficiency.", "texts": [ " To increase productivity, a typical Gaussian laser beam was shaped into a square flat-top beam. Based on the numerical simulation of the fluid flow of the powder, a scanning nozzle was developed by optimizing the design conditions; this improved the deposition efficiency and quality of the metal powder. Finally, a single bead and multiple layers were fabricated by assembling the nozzle onto the deposition head and fitting it to the laser equipment. The conceptual design of scanning- directed energy deposition (SDED) for the high-throughput DED method is shown in Fig. 1(A). In general, laser-based AM uses a circular beam of several millimeters, in the form of a Gaussian distribution; this is mainly used for precision deposition. The laser beam irradiates the metal powder (on the surface of the base material) and melts it, thereby forming metal beads. Therefore, the width of the metal bead depends on the spot size of the laser beam irradiating the surface of the base material. Using our proposed method, we could deposit a large-width metal bead by utilizing a square flat-top 8 mm \u00d7 3 mm beam. In the metal AM processes, several beads are overlaid layer by layer to form the desired shape. Therefore, the waviness of the bead surface should be minimized to achieve precise and defect-free deposition. Fig. 1 (B) and (C) show the waviness of the metal beads formed on the surface of the base material, according to the number of outlets through which the metal powder is sprayed. In general, when a metal bead is formed, the center of the bead has a large height compared to that of its edge. Due to the process of layer-by-layer metal deposition, as the height difference (\u0394H) increases, pores and defects may occur between each layer. Therefore, it is assumed that as the number of outlet channels (for metal powder spray) increases, the \u0394H decreases resulting in a decrease in waviness due to the decrease in fluid flow rate. Fig. 1(D) and (E) show the area of the metal powder sprayed with respect to the distance between the metal powder nozzle and the base material surface. The metal powder sprayed from the nozzle gets deflected to the sides, while traveling to the surface of the base material, which decreases the deposition efficiency. Therefore, to improve the efficiency, we considered decreasing the distance between the base material surface and the nozzle (d\u2192d\u2019). This could make the powder area and the laser beam area irradiated on the base material surface approximately equal (AP \u2248 AL) Based on these considerations, we developed a high-throughput DED method with an optimized scanning nozzle to spray powder using a square laser beam" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.16-1.png", "caption": "Fig. 3.16 Aerial Gun, tracking a target at p = (x/y/z)", "texts": [ " Angles from rotation sequences that involve all three axes (x-y-z, y-z-x, z-x-y, x-z-y, z-y-x, y-x-z) can be either called \u201cTait\u2013Bryan angles\u201d, in honor of the Scottish mathematical physicist Peter Tait (1832\u20131901), who was\u2014 together with Hamilton\u2014the leading exponent of quaternions, and the Welshman George Bryan (1864\u20131928), the originator of the equations of airplane motion or \u201cCardan angles\u201d, after the Renaissance mathematician, physician, astrologer, and gambler Jerome Cardan (1501\u20131576), who first described the cardan joint which can transmit rotary motion. And angles that have the same axis for the first and the last rotation (like the Euler sequence above) are called \u201cproper Euler angles\u201d (z-x-z, x-y-x, y-z-y, z-y-z, x-z-x, y-x-y). Consider tracking a helicopter flying from the horizon toward an aerial gun, as indicated in Fig. 3.16. The helicopter flies toward the gun site and is tracked by the gun in elevation ( ) and azimuth ( ). When the helicopter is immediately above the gun site, the aerial gun is in the orientation indicated in Fig. 3.14. If the helicopter now changes direction and flies at 90\u25e6 to its previous course, the gun cannot track this maneuver without a discontinuous jump in one or both of the gimbal orientations. There is no continuous motion that allows it to follow the target\u2014it is in \u201cgimbal lock\u201d. Note that even if the helicopter does not pass through the gimbal\u2019s zenith, but only near it, so that gimbal lock does not occur, the system must still move exceptionally rapidly to track the helicopter if it changed direction, as it rapidly passes from one bearing to the other", " Thereby, instructions can be executed in parallel and can be dramatically optimized. An important application of homogeneous coordinates are the \"Denavit Hartenberg parameters\". These are used in mechanical engineering to denote position and orientation of an end-link in robot manipulators (see Appendix B). 48 3 Rotation Matrices A few examples of the application of rotation matrices may help to show how to use them in practical applications. An aerial gun is mounted like a nautical gimbal: the outermost rotation is always about an earth-vertical axis (Fig. 3.16). In the starting orientation (\u03b8/\u03c6 = 0/0), the barrel of the aerial gun points straight ahead, i.e., along bx. Task: When a target appears at p = (x, y, z), we want to reorient the gun such that the rotated gun barrel, which after the rotation points in the direction of b\u2032 x, points at the target. Solution: Taking the rotation matrix in the nautical sequence (Eq. 3.24), with \u03c8 = 0 since a rotation about the line of the gun barrel is not relevant, we get Rnautical(\u03c8N = 0) = [ b\u2032 x b \u2032 y b \u2032 z ] = \u23a1 \u23a3 cos \u03b8N cos\u03c6N \u2212 sin \u03b8N cos \u03b8N sin \u03c6N N sin \u03b8N cos\u03c6N cos \u03b8N sin \u03b8N sin \u03c6N \u2212 sin \u03c6N 0 cos\u03c6N \u23a4 \u23a6 . With the first column b\u2032 x = p |p| , this leads to \u03c6N = \u2212 arcsin ( pz\u221a px 2+py 2+Pz 2 ) \u03b8N = arcsin ( py\u221a px 2+py 2+pz 2 \u00b7 1 cos\u03c6N ) . (3.35) 3.6 Applications 49 Note that a combination of a horizontal and a vertical rotation of the object in a well-defined sequence uniquely characterizes the direction of the forward direction. With eye movements, this is the line of sight, or gaze direction; with a gun turret on a ship this is the direction of the gun barrel (see Fig. 3.16). However, this does not completely determine the three-dimensional orientation of the object, since the rotation about the forward direction is still unspecified. A third rotation is needed to completely determine the orientation of the object. This third rotation\u03c8 would not affect the direction in which the gun is pointing, it would only rotate around the pointing vector. For a quaternion solution to the targeting problem, see also Sect. 4.5.1. Another frequent paradigm is a projection onto a flat surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001385_1.4029766-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001385_1.4029766-Figure1-1.png", "caption": "Fig. 1 Schematic diagram illustrating the autogenous laser brazing setup. The wires are rotated, while the laser simultaneously scans toward the interface, but is turned off before crossing over to the SS side. The angle h corresponds to the apex angle of the cup/cone configuration.", "texts": [ " Nickel titanium is nearly equiatomic, and iron is the most prevalent element in SS. A Ti\u2013Ni\u2013Fe ternary phase diagram can thus be used to determine which phases are formed when these three elements are melted and brought into contact with each other [11]. All of the potential phases that form (NiTi2, FeTi, Fe2Ti, and Ni3Ti) are brittle intermetallics. Excessive amounts of these brittle phases will decrease the fracture strength of the weld. Using lasers as the input heat source can help to limit the formation of these phases by preventing overmelting. Figure 1 shows a schematic diagram of the autogenous laser brazing setup. The laser starts a fixed distance away and is scanned at a constant velocity toward the interface. Simultaneously, the wires are rotated, to ensure radially uniform heating. Thermal accumulation occurs at the interface since there is a thermal resistance between the two wires. The thermal accumulation results in a high localization of heating, so that only a very limited amount of the NiTi will melt and then wet onto the SS wire. Thus, a narrower region of intermetallics forms, resulting in increased fracture strengths" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.1-1.png", "caption": "Fig. 3.1 Similarity between one-dimensional translations (left) and single-axis rotations (right): Both require the selection of a reference, and both are characterized by a single parameter", "texts": [ " After choosing an arbitrary point in space as the reference position, the position of each point is defined by three translations away from the coordinate center, e.g., forward, left, and up. Here, it is worth pointing out a seemingly obvious fact: the final location of the object is independent of the sequence of these translations. If we move first 10m right and then 15m forward, we end up in the same location as if we had moved first 15m forward, and then 10m right. This property is referred to as the commutativity of translations (see also Fig. 5.1). The description of orientation is done in a similar way (Fig. 3.1). First, an arbitrarily chosen orientation is defined as reference orientation. Once that is done, any other orientation can be described by three parameters: an object can not only be 1An inertial frame is a frame of reference in which a body remains at rest or moves with a constant linear velocity unless acted upon by forces. An inertial reference frame does not have a single, universal coordinate system attached to it: positional values in an inertial frame can be expressed in any convenient coordinate system", "2 Rotations in a Plane 33 As a result, a rotation of a 2-D-vector, expressed as a complex number c, by an angle \u03c6, can be written as c\u2032 = e j\u03c6 \u2217 c = e j\u03c6 \u2217 (r \u2217 e j\u03b8 ) = r \u2217 e j (\u03c6+\u03b8). (3.9) Note: In mathematics and physics, the square root of \u22121 is typically denoted with i , whereas in many technical areas j is used. In both Python and Matlab, j can be used: x = -1+0j np.sqrt(x) >>> 1j In polar coordinates, the similarity between one-dimensional translations and singleaxis rotations becomes obvious (Fig. 3.1). Task: If a gun originally pointing straight ahead along the +x axis is to shoot at a target at P = (x, y), by which amount does the gun have to rotate to point at that target (Fig. 3.5)? Solution: The gun barrel originally points straight ahead, so the direction of the bullet aligns with nx. The rotation of the gun is described by the rotationmatrixR = [ cos \u03b8 \u2212 sin \u03b8 sin \u03b8 cos \u03b8 ] =[ n\u2032 x n \u2032 y ] . The direction of the gun barrel after the rotation is given by n\u2032 x = p |p| , which is also the first column of the rotation matrix R" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure13-1.png", "caption": "Fig. 13. Finite element model of gearbox.", "texts": [ " The model is simplified as follows: the unimportant parts of the box, such as bolt hole, chamfering, and oil sight glass, are ignored; the bearing is simplified as the inner and outer ring models; the outer ring is fixed with the box; the inner ring is fixed with the shaft. The information about the finite element model of the gearbox is shown in Table 4. Dynamic meshing forces are applied on the meshing line of each gear pair as dynamic excitation of the system. The finite element model of the gearbox is shown in Fig. 13, and the directions of X, Y, and Z in the diagram refer to the transverse, vertical, and axial directions, respectively. In this study, the rated input speed and input power of the gearbox and the dynamic meshing forces of the each gear pair are applied as the internal excitation to calculate the dynamic response. In the calculation, the input speed is 1500 r/min, the input power is 45 kW, and the eccentricity errors of gear 1 are 0 and 20 \u03bcm. Dynamic response of the gearbox is calculated with the mode superposition method in ANSYS software" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001020_s00170-015-7033-2-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001020_s00170-015-7033-2-Figure5-1.png", "caption": "Fig. 5 Coordinate of the grinding machine", "texts": [ " The spiral groove of the rake face is formed by sweeping the grinding wheel along the cutting edge curve. To obtain desired edge shape, accurate grinding trajectory should be proposed based on the mathematical model of the double-circular-arc cutting edge. Grinding simulation is required to predict the grinding process and optimize the fabrication due to complexity of the CNC grinding of the milling cutter. The cutting edge design and grinding trajectory are tested by grinding simulation in this work. As shown in Fig. 5, the NC grinding machine has five axes, that is, linear motion axes X, Y, Z and rotary axes A and B. The grinding coordinates, including the workpiece coordinate, normal-section local coordinate, grinding wheel coordinate, and machine tool coordinate, are shown in Fig. 6. The workpiece coordinate [OtXtYtZt] attaches on the workpiece to describe the geometric shape of the helix cutting edge of the workpiece. The origin of the normal-section local coordinate [OpXpYpZp] locates on the helix edge curve to determine the position of the grinding wheel to the machining site and solve the tool position parameters of the machining surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure15.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure15.7-1.png", "caption": "Figure 15.7 Forming a drill guard", "texts": [ " Shapes other than simple bends can be carried out by heating the complete piece of material in an oven. To avoid marking the surface, the material can be placed on a piece of brown paper. The time in the oven depends on the type of material and its thickness, and time must be allowed for the material to reach an even temperature throughout. Acrylic sheet material is easily worked at 170 \u00b0C, 6 mm thickness about 30 minutes in the oven. Again, a simple former can be used to obtain the required shape, e.g. in making a guard for a drilling machine, Fig.\u00a015.7. Drape forming (also called oven forming) is carried out commercially on small and large parts where the plastics material is pre-heated in an oven to the forming temperature, placed in a mould and held in place while it cools. Applications include D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 15 Plastics 15 245 of gases and easy removal of the product after moulding. Moulding tools are manufactured from tool steel, hardened and tempered to give strength, toughness and a good hard-wearing surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000345_j.optlastec.2019.105723-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000345_j.optlastec.2019.105723-Figure2-1.png", "caption": "Fig. 2. Scheme of the 3D FG sample fabrication based on the NiCrSiB alloy matrix with WC addition via the SLM process.", "texts": [ "07 \u03bcm, power up to 100W) operating in a continuous mode independently from each other; own deflectors for each wavelength to implement the laser influence (LI) scanning over the powder surface; soft process control via a personal computer; interchangeable focusing lenses; mechanism for delivering and leveling several types of powder mixtures simultaneously; cylindrical platform moving in the vertical direction, on which 3D parts are layerwise fabricated and in-situ diagnostics of the SLM process. The procedure for the FG structures fabrication and gradient 3D parts based on the MMC with the NiCrSiB alloy matrix is shown in Fig. 2, we developed it earlier [11,34,38]. The hatching space was equal to the laser beam diameter db \u2013 70 \u03bcm and layer depth H was ~0.2mm. The argon atmosphere was created by replacing air from the hermetic synthesis chamber before each 10-layer cycle. The layers were prepared of premixed NiCrSiB+WC powders on a substrate by the following scheme (Fig. 2): after probed NiCrSiB laser cladding, the first ten layers were of NiCrSiB with 5%vol. WC, the second ten consisted of 90% NiCrSiB+ 10%vol. WC, and the third layers - of 85% NiCrSiB+15%vol. WC. The laser scan rate V ranged from 5 to 20 cm/s, laser power P - from 20 to 90W. Two regimes of manufacturing have been studied \u2013 with the substrate and chamber additional heating up to 300 \u00b0C and without it. Each second layer was formed on the bottom layer after turning it by 90\u00b0 (L \u2013 longitudinal, T - transversal in the Fig. 2). After the etching, cross sections of multi-layered melting samples were subjected to metallurgical analysis using the optical microscope (Neophot 30M, Carl Zeiss) equipped with a digital camera. The etching was carried out on our samples in the next test solution: 7.5 ml of HCl+ 2.5 g of FeCl3+ 100ml of H2O. The 3D samples obtained under the optimized regimes were analyzed by microhardness tester PMT-3M (OKB SPECTR Ltd., St. Petersburg, Russia). Microstructures were studied by scan electron microscopy LEO 1450 (Carl Zeiss Company) equipped with an energy-dispersive X-ray microanalyzer (INCA Energy 300, Oxford Instruments)", " One of the recommended layerwise SLM regime was regime the P=90W, v= 20 cm/s with the hatching space equal to the laser beam diameter and layer depth H~0.2mm that prove to be acceptable for three variants of the WC inclusions (5\u201310\u201315% vol) in the NiCrSiB matrix. Fig. 3 presents the visual form of a 3D cube from the MMC based on the NiCrSiB alloy in which the content of the ceramic doping additive equal to approximately 5% by volume at the base and 10% by vol. in the middle, can further increase towards top up to 15% by vol. according to the scheme shown in (Fig. 2). It is clear that a significant roughness of the fused surface structure without additional heating of substrate and powder in the chamber. Based on the results of optimizing the SLM regimes for self-fluxing nickel-chrome alloys, we concluded that there is no alternative to additional heating in the SLM chamber up to 500\u2013700 \u00b0C, as it was proposed earlier in [38]. The results of optical microscopy (OM) of remelted structures after the SLM in the materials are shown in Fig. 4. Fig. 4a shows the microstructures after the SLM of the pure NiCrSiB alloy without substrate heating" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000447_j.snb.2019.127152-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000447_j.snb.2019.127152-Figure1-1.png", "caption": "Fig. 1. Construction of the b-BAgmE: the inner silver core scheme (A), the electrode construction scheme (B), photography of the sensor (C). Ag foil piece (a), Ag wire (b), flattened Ag wire end (c), epoxy resin (d).", "texts": [ " All solutions were prepared using 4 times distilled water (quartz). The b-BAgmE was prepared from a silver foil (Alfa Aesar, Germany, 99.95%, 25 \u03bcm), a silver wire (The Mint of Poland, Poland, 93%, \u03d5=0.7mm) and TRANSLUX D180 resin (AKSON, France). For polishing the electrode surface, MicroPolish\u2122 Alumina suspensions (1.0, 0.3 and 0.05 \u03bcm, Buehler Micropolish II, Lake Bluff, IL USA) and MasterTex polishing cloth (Buehler), were used. A schematic diagram of the b-BAgmE fabrication steps and its photography are illustrated in Fig. 1. A 3D animation of the electrode construction is available in the electronic version of this article (Animation 1 ). The metallic core of the sensor (Fig. 1A) was prepared with a silver wire (\u03d5=0.7mm) and a silver foil (25 \u03bcm). One end of the Ag wire was flattened with a vise over a length of ca. 20mm. At a distance of ca. 10mm from the end of the wire it was bent into the shape of the extended letter \u201cU\u201d. In the slit formed in this manner, a properly cut piece of Ag foil (4.50 x 12mm) was symmetrically placed. Afterwards, sides of the Ag wire were mechanically tightened to stabilize the construction and to ensure electric contact. Prepared in this way main element of the electrode was cleaned with acetone and covered with epoxy resin" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003668_j.egyr.2021.05.047-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003668_j.egyr.2021.05.047-Figure4-1.png", "caption": "Fig. 4. (a) Schematic diagram of wind-driven nanogenerator based on the as-prepared rTENG. (b) short-circuit current of rTENG under various wind speeds.", "texts": [ " In order to investigate the universality f the as-developed GO/PVDF film-based rTENG in harvesting arious frequency mechanical energy, the relationships between he electric output signals and rotation speed were systematically nvestigated. As shown in Fig. 3c and d, both output current and utput voltage increased with the rotation speed increasing. In practical use, a wind turbine nanogenerator was designed ased on the as-prepared rTENG as schematically shown in ig. 4a. The short-circuit current were measured under various imulated wind speeds and present a certain functional relationhip (Fig. 4b&S5). We should note that the relationship between rpm and the speed of wind can be obtained by Equation S1 in supporting information. For the as-prepared 0.25 m% GO@PVDF film-based rTENG, increasing wind speed from 2.8 m/s to 12.1 m/s resulted in significant improvement of current from 4 \u00b5A to 12.5 \u00b5A. In summary, a flexible and eco-friendly rTENG based on GOfilled PVDF film and EC layers was developed. Moreover, the concentration of GO chemically deposited on the surface of the PVDF layer revealed a significant effect on the performance of the rTENG" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002528_s0263574720000429-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002528_s0263574720000429-Figure1-1.png", "caption": "Fig. 1. The proposed soft finger with visual sensor. (a) Structure of the proposed soft finger. (b) Prototype.", "texts": [ " Besides, the inner chamber is in close contact with the outer chamber to prevent expansion cracking. When the soft finger is bending, the camera embedded into the sealing device is used to capture the images of the deformation of the inner chamber. The camera\u2019s resolution is 640 \u00d7 480 and the frame rate is as high as 30 fps. Referring to human\u2019s finger, the length of inner chamber is selected as 10 cm. And the length of each color region is divided as follows: 4 cm for red region, 3 cm for blue region, and 3 cm for yellow region. The proposed soft finger is shown in Fig. 1. In this section, the perception algorithm based on image preprocessing and deep learning is proposed. We extract boundary of color regions from inner chamber image and get position of marker dots from label image by image preprocessing. The bending state of finger is predicted from boundary by convolutional neural network (CNN). The proposal of the algorithm is shown in Fig. 2. https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574720000429 Downloaded from https://www.cambridge" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.2-1.png", "caption": "FIGURE 2.2 Schematic layout tire structure.", "texts": [ " Changes in tire load will change the tire performance, which must be accounted for in the vehicle handling analysis. When the driver is cornering, the outer tires are loaded and the inner tires are unloaded. When the driver is braking or accelerating, the tire load shifts between the front and rear wheels. An increase in vehicle speed will in general lead to more critical adverse tire road conditions. All these effects depend on the tire inner pressure. We will take a closer look at the structure of the radial tire (Figure 2.2). The term \u201cradial tire\u201d refers to the radial plies, running from bead to bead, with the bead being the reinforced (with an embedded steel wire) part of the tire, connecting the tire to the rim. However, radial plies do not give the tire sufficient rigidity to fulfill the required performance under braking and cornering conditions. For that reason, the tire is surrounded by a belt with cords (steel, polyester, Kevlar, etc.) that are oriented close to the direction of travel. The radial plies give good vertical flexibility and therefore, good ride comfort (in case of road irregularities)", " Because of the structural differences between both tires, the tread motion is reduced for the radial tire compared to the cross-ply tire, which also contributes to better fuel economy (reduced rolling resistance, see also Section 2.3). It has been shown by Moore [29] that bias-ply tires show significantly higher concentrations of shear stress, as well as normal contact pressure, at the shoulders of the tire, compared to radial tires. The main contact between tire and road is through the tread area. Figure 2.2 indicates a tread pattern that is shaped to channel water away, with straight and s-shaped grooves that move from center of the tire to the side. We also indicated very small cuts in the pattern, referred to as sipes. These sipes are typical for winter tires and allow small motion between tread elements for rolling tires, leading to effectively larger friction on icy and snowy surfaces. In the next section, we begin with a description of the input and output quantities of a tire. Determining what forces and moments are acting on a tire, and what input variables (such as slip, camber, and speed) these forces and moments depend on defines our language to define tire characteristics", " Under wet road conditions, the longitudinal force coefficient maximum level drops, to levels on the order of 0.6 0.8 for a wet road, to 0.4 0.5 for snow, and to levels of 0.2 0.4 for ice. A special case is given if a significant amount of water is present on the road. To maintain contact between tire and road, the water must be evacuated. This property may be improved by adjusting the tread block pattern of the tire (longitudinal grooves, or grooves curved in an outward direction guiding the water in a radial direction away from the tire, see also Figure 2.2). With increasing speed, there is less time to remove the water and the contact zone is further reduced. Consequently, the brake force and therefore, the friction coefficient drops significantly with vehicle speed. At a certain speed, the tire may float entirely on a film of water (hydroplaning), and the friction coefficient drops to very low values (,0.1). In other words, hydroplaning occurs when a tire is lifted from the road by a layer of water trapped in front of and under a tire. One usually distinguishes between dynamic hydroplaning (water is not removed fast enough to prevent loss of contact) and viscous aquaplaning (the road is contaminated with dirt, oil, grease, leaves, etc" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002914_978-3-030-60330-4-Figure7.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002914_978-3-030-60330-4-Figure7.7-1.png", "caption": "Fig. 7.7 Self-modulation of a quasiharmonic wave (a) and evolution of its spectrum (b)", "texts": [ " It is known that under certain conditions a quasiharmonic wave is unstable with respect to breaking up into individual wave packets (this effect is called a modulation instability or a self-modulation) [25]. The presence of such instability in a system is determined by the Lighthill criterion [26]: dvgr dk \u03c3 < 0. (7.39) Speaking in spectral language, themodulation effect is characterized by increasing of side components in the modulated wave spectrum. The energy will be pumped into these components from the central part of the spectrum of perturbations. Figure 7.7 schematically shows the self-modulation process of a quasiharmonic wave (a) and evolution of its spectrum (b). 7.3 Self-modulation of Shear Strain Waves \u2026 161 In the case at issue, as follows from Eqs. (7.37) and (7.38), dvgr dk \u03c3 = \u2212 6b1b2 \u03b5 k2 = \u2212 9\u03b3 2R2(c23\u2212c\u030322) 16\u03b5(c21\u2212c\u030322) c\u0303 2 2 k2. Since from Eqs. (7.26) and (7.20a), it is visible that c\u030322 < c23 < c21, then, according to the Lighthill criterion, the modulation instability will be observed in the considered medium for all the allowable values of the microstructure parameters [20]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001910_978-3-319-15010-9-Figure9.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001910_978-3-319-15010-9-Figure9.2-1.png", "caption": "Fig. 9.2 (a) Graph D and (b) resultant node and edge weights importance pertaining to Examples 9.3 and 9.6", "texts": [ " As each of these input sets satisfies the condition in Theorem 9.1, the pairs .A;B.f1g//, .A;B.f2g//, and .A;B.f3g// are weakly s-controllable. 154 9 Security and Infiltration of Networks: A Structural Controllability. . . The following example examines a \u201cHighjack and Eavesdrop\u201d scenario for an unknown realization of Model (8.1). Theorem 9.1 is applied to acquire the most vulnerable weakly controllable attack vectors and reason about the relative node security in the network. Example 9.3 (Highjack and Eavesdrop). Consider the graph D in Fig. 9.2a and network dynamics corresponding to a realization of .A.D/;B.S/;C .S// in Model (8.1). A weakly controllable inputs set S presents a particularly attractive attack vector for an infiltrator. The node set S almost always provides a controllable input and observable output sets to effectively control (highjack) the network and/or monitor (eavesdrop on) the network through system identification. Applying Theorem 9.1 to all possible input sets the smallest cardinality attack vectors involve the successful infiltration of two nodes", " Specifically, the nine smallest attack vectors are of the form fj; 7g and fi; 8g, where j 2 f1; 3; 4; 5g and i 2 f1; 2; 3; 5; 6; 7g. All larger attack vectors contain one of these pairs. In aggregate there are 821 attack vectors, the 345 attacks involving at most four nodes are distributed as A D f38; 36; 38; 37; 38; 36; 60; 62g where the number of attacks involving node i is A.i/. From a network design perspective the most to least vulnerable node sets are f8g ; f7g ; f1; 3; 5g ; f4g ; f2; 6g, providing a priority ordering for security. Figure 9.2b indicates this ordering. Further, if nodes 7 and 8 are completely secured against attachment, then there will be no input set that will render the system controllable or observable. An attractive feature of s-controllability is the provided controllability guarantees in the face of perturbations. Specifically, as long as the interconnections in the graph remain intact and no new ones are added controllability will be maintained. One can consider this a type of controllability robustness\u2014an often elusive feature of system dynamics, with progress made in the area of controllability of interval matrices [18]", " The matchings associated with input set f2g appear in Fig. 9.3. We now apply Theorem 9.4 to a \u201cDisrupt\u201d scenario where an infiltrator perturbs edge weights, potentially removing them so as to reduce the effectiveness of the control input into network running a realization of Model (8.1). Examining the effect of edge failures on strong s-controllability of the input set one can identify the critical edges in the network and identify the most significant security vulnerabilities. Example 9.6 (Disrupt). Consider the graph D in Fig. 9.2a and network dynamics corresponding to a realization of .A.D/;B.S/;C .S// in Model (8.1), where S D f3; 5; 7g. Applying Theorem 9.4 reveals that the system is strongly s-controllable and observable. If an infiltrator\u2019s objective is to disrupt the effective control and monitoring of the network, a viable strategy is to perturb the network\u2019s interconnection strengths, i.e., its edge weights. As long as no interconnections are References 157 broken, input set S can still control and monitor the network", "3, where if S1 is weak and S1 S2 then S2 is weak, the same condition does not hold for edge failure attack vectors. For example, the removal of E2 renders the system no longer strong s-controllable but also removing 1 ! 1 returns strong s-controllability. Examining all 1,208 successful attacks involving three edges, there were 106, 89, 67, and 60 successful attacks involving each of the nodes in E1, E2, E3, and E4, respectively, whereE3 D En nS3 iD1 Ei o andE4 D f1 ! 1; 5 ! 6g. Consequently, the network edge sets E1; : : : ; E4 present a preferential ordering of the distribution of security resources. Figure 9.2b depicts this ordering. This chapter presents an analysis of the security of networked system topologies using weak and strong s-controllability. Focusing on \u201cDisrupt\u201d and \u201cHighjack and eavesdrop\u201d attack scenarios, we propose controllability metrics to identify vulnerable nodes and critical edges of a network. This was accomplished through a computationally efficient matching condition on weak and strong s-controllability. Future work of particular interest involves establishing conditions for output weak and strong s-controllability and their implications for network security" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002092_0954406216682768-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002092_0954406216682768-Figure10-1.png", "caption": "Figure 10. Illustration of the designed preload-adjustable mechanism.", "texts": [ " The structure of the loading mechanism. is installed to measure the tension and compression. In this study, in order to ensure that the loading mechanisms just apply axial force on the working table without extra torque, it needs to guarantee that the compressed length of springs is the same by controlling the motors synchronously. A preload-adjustable double-nut mechanism that can adjust preloads by regulating the distance between the master nut and slave nut, and measure preloads in real time is designed as presented in Figure 10. Regulating the preload is executed by rotating the adjustment nut. In order to reduce the friction when it is rotated, a thrust bearing is installed adjoining to the adjustment nut. Two disc springs are installed between two nuts. When the adjustment nut is rotated, the disc springs are compressed and axial force is exerted on the master nut and slave nut. The distance between the master nut and slave nut becomes larger at the same time, and the balls inside the master nut and slave nut are squeezed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003481_s00170-021-07005-6-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003481_s00170-021-07005-6-Figure11-1.png", "caption": "Fig. 11 Predicted results of the longitudinal stress (\u03c3xx) at different process times in a and c and longitudinal displacement (Ux) at different process times in b, and d", "texts": [ " In LMD process, the laser beam moves according to the path plan and the cyclic scanning sequence causes cyclic temperature change at each monitoring location. It shows ten temperature peaks for conventional method and five temperature peaks for simplified method. The conventional method shows temperature evolution with higher resolution in each individual deposited layer. Except that, the overall trend of temperature evolution, i.e., increasing temperature peaks and temperature difference between two consecutive cycles, are maintained for both simulation methods. Figure 11 shows the longitudinal stress and displacement distribution in x-direction at process times of all deposition done and cools to room temperature in 2-layer by 2-layer method. At the end of the deposition process in Fig. 11a, tensile stress accumulates near the top free end and large compressive stress builds up near the substrate-deposit interface, reach the values of 188 MPa and 984 MPa, respectively. This trend is similar to the predicted results by conventional method reported in Fig. 8 and has the same magnitude level. When the coupon cools down to the room temperature, the largest residual stress with the value of 473MPa was observed in Fig. 11c, which is a little bit higher than that in Fig. 8i due to the assumption of activating two layers each time. The tensile stress within the top deposited layer is caused by the contraction of the molten material after cooling. For layers underneath the top layer, tensile stress is reduced and gradually changes into compressive stress because of the annealing effect by subsequent deposition layers. The longitudinal displacement profile (along x direction) in Fig. 11b and d consists of large values near the top and two ends of the as-built part, while smaller displacement occurs in the center of the part. Same level of displacement compared with conventional model can be found here. In addition, residual stress field and total displacement field from simplified method is presented in Fig. 12, as a comparison to the conventional method in Fig. 9. Similarly, even if the difference exists for simplified method, the overall average stiffness of the structure does not show much change" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure3-1.png", "caption": "Fig. 3. 3D printed axial node and bending node.", "texts": [ " In the Binder Jet method, a binder is selectively deposited onto the powder bed, bonded metal particles all together to form a fragile and porous part one layer at a time. The process of separating the powder from the 3D printed porous model (by using shaking table) is called de-powdering step. Then a solid and strong part will be made by flowing bronze into the pores in an infiltration process [20]. To avoid breaking the porous model during depowdering step, six auxiliary members are added to the model to be 3D printed and cut off before test. The 3D printed nodes and auxiliary members are shown in Fig. 3. To obtain the properties of the printed stainless steel, a series of tensile tests are carried out on six 3D printed dog-bone samples and the test results are shown in Fig. 4. As can be seen, the tested modulus of elasticity, strength and elongation for the 3D printed material vary in different extent. This uncertainty in material properties may cause uncertain results in the structural tests. For numerical simulation purpose, the average of the test data is employed and simplified by a multilinear curve" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002701_aero47225.2020.9172614-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002701_aero47225.2020.9172614-Figure16-1.png", "caption": "Figure 16: A model of a next iteration of the Tetrahedral rotorcraft with three rotors configured to produce force vectors pointing inward while the remaining rotor produces a force vector pointing inward.", "texts": [], "surrounding_texts": [ "The purpose of this research is to set the foundation for future models of the Tetracopter as well as the Fractal Tetrahedron Assembly. A prototype of the Fractal Tetrahedron Assembly with multiple Tetracopters will be built and the precise control laws at different level of the assembly will be studied. Although this research proves that the rotor configuration of Authorized licensed use limited to: Middlesex University. Downloaded on September 02,2020 at 09:13:47 UTC from IEEE Xplore. Restrictions apply. 11the Tetracopter is feasible, there are several parameters of the current design that could increase the efficiency and control of the rotorcraft. One of the main improvements, highlighted by the CFD results, is the aerodynamics of the frame and the relative positioning of the rotors. In the future, the wake interaction between the top propeller of the Tetracopter and the bottom ones will be studied in more details. Reducing the height of the Tetracopter could potentially improve its efficiency, at the cost of a reducuded rigidity when assembled. Future research will also include more sophisticated hardware for the Tetracopter prototype in order to achieve the goal of in-flight self-assembly. Other versions of the Tetrahedral rotorcraft using a single tetrahedron with a propeller on each face will also be designed. The relationship between the minimum number of motor failures needed to hinder a FTA from hovering and the number of rotors in the FTA is also of interest for future works." ] }, { "image_filename": "designv11_22_0002902_0954406220967693-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002902_0954406220967693-Figure5-1.png", "caption": "Figure 5. Lumped parameter of the two-stage straight bevel gear system.", "texts": [ " Dynamic equations governing the motion of the system The two-stage straight bevel gear contains three block. Each block consists of two wheels (ji) and shaft (j). i denotes the index of the wheel in block (j) as shown in Figure 4. The bearing supporting the blocks are placed at the middle of the shafts, and they are off the geometric center of the wheels by LA j . The frame E 1\u00f0 \u00de and E 3\u00f0 \u00de are placed at the geometric centers of the gear (12) and gear (31) respectively. For the case of the block (2), the frame E 2\u00f0 \u00de is placed at the middle of the shaft (2). Figure 5 shows the lumped parameter of the twostage straight bevel gear. The translational coordinates of each block (j) xj, yj, zj are assigned to translate along E j\u00f0 \u00dex; E j\u00f0 \u00dey; E j\u00f0 \u00dez respectively. The angular coordinates of the block (j) hji, wj, /j are assigned to rotate about E j\u00f0 \u00dex; E j\u00f0 \u00dey; E j\u00f0 \u00dez respectively. The bearings using in this system can be modeled as diagonal stiffness matrices Kbearing;j \u00bc diag 0 k w j k/j kxj k y j kzj h i (10) The block (j) are linked with the block (j\u00fe1) by spring acting along the line of action and perpendicular to the tooth surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000950_icsns.2015.7292433-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000950_icsns.2015.7292433-Figure1-1.png", "caption": "Fig. 1 (a-b) Typical Liquid Cryogenic Cylinders", "texts": [ " This is very critical in the case of delayed launching of aerospace vehicles after filling the cryogenic fuel in the rocket tanks. Accurate fluid level estimation can help the mission director to take a thoughtful decision for refilling the tank before its launch. (a) Typical cryogenic tank Note that the refilling time of cryogenic tank might lead to an altered launch window and other trajectory prediction complications. Therefore an accurate measurement of the cryogenic fluid level is inevitable for any mission. Figure 1 shows the illustration of cryogenic tanks, gauge and its valve configurations. The liquid level control systems are widely used for monitoring of liquid levels, reservoirs, silos, and dams etc. But cryogenic sensors are very critical in design. Usually, liquid level control kind of systems provides visual multi level as well as continuous level indication. Audio visual alarms at desired levels and automatic control of pumps based on user\u2019s requirements can be included in this management system. Proper monitoring is needed to ensure water sustainability is actually being reached, with disbursement linked to sensing and automation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003429_j.tws.2021.107585-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003429_j.tws.2021.107585-Figure1-1.png", "caption": "Fig. 1. Dished ends according to EN 13445; (a) - torispherical, (b) - ellipsoidal, (c) - hemispherical.", "texts": [ " The Ritz method has not been applied to the practical cases of the pressure vessels. The analytical solution for such problems can be imprecise, therefore the use of Ritz method may allow to obtain satisfactory results in the analyses of stress and deformations in the standard pressure vessels. Additionally, the semi-analytical formulation of the problem can include orthotropic properties of the material. 2. Geometry of standard dished ends The standard dished ends according to EN 13445 [4] are ellipsoidal, torispherical and hemispherical (Fig. 1). The latter type is not considered within the paper as the accurate solution for those can be obtained in analytical approach with the use of the edge effect theory. The shape of torispherical ends is based on two radii i.e. \ud835\udc5f\ud835\udc56 \u2013 inside radius of curvature of a knuckle and \ud835\udc45\ud835\udc56 \u2013 inside spherical radius of the central part of the torispherical end. According to EN 13445 [4], ellipsoidal end shall be designed as nominally equivalent torispherical ends with \ud835\udc5f\ud835\udc56 = \ud835\udc37\ud835\udc56 ( 0.5 \ud835\udc3e \u2212 0.08 ) , (1) and \ud835\udc45\ud835\udc56 = \ud835\udc37\ud835\udc56 (0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003510_s12206-021-0424-4-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003510_s12206-021-0424-4-Figure4-1.png", "caption": "Fig. 4. Deployment error indication during membrane deploying process.", "texts": [ " In addition, the flatness error refers to the distance between two parallel planes containing the actual plane or any specified range of the actual plane, which must meet the minimum condition [17]. Based on this, the neutral surface of the membrane during the deploying process is analyzed. Two planes parallel to the ideal membrane fully deployed plane are used to contain the actual deployment neutral surface. The minimum distance between the two parallel planes is taken as the deployment error. The deployment error is indicated during membrane deploying process, as shown in Fig. 4. The deployment rate and deployment error are used as the evaluation indexes of the membrane deployment results. Due to the elastic-plastic deformation during the membrane deploying process, although the crease is not fully deployed when F = 1.46 N/mm, the deployment rate of the membrane has reached 100 %. At this time, the deployment error is 0.17 mm, and the deployment angle is about for 82\u00b0. When the membrane thickness t and the folding height h are both constant, the F-\u03a6 relationship of Z-folded membranes of different sizes may be different when deployed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002408_s11012-020-01162-w-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002408_s11012-020-01162-w-Figure1-1.png", "caption": "Fig. 1 The schematic diagram of the power system of the presented flapping wing", "texts": [ " The impact of any design variables or control parameters on the bird performance can be easily studied through this integrated and systematic model. The under investigation flapping ornithopter in this study includes the main body, the two-section wings with the span of 1 m, a pair of DC motors and power transmission systems. In this section, the governing equations of the main components of the power system (propulsion and power transmission systems), twosection wings and the main body, which are modeled with the help of the bond graph method, are presented. 2.1 Power system Figure 1 illustrates the schematic diagram of the presented bird power system, including the propulsion system and the power transmission mechanism. In this system, a pair of DC motors is used as the propulsion system. The motor armature consists of an inductance and a resistance that converts the electrical energy to the mechanical one. The motor characteristics include the armature voltage Ea, the nominal voltage Vt, the armature current Ia, the armature resistance ra and the armature inductance La. The governing differential equations of this model are as follows: d dt Ia \u00bc ra La Ia Kv La xa \u00fe Vt La d dt xa \u00bc Kt J Ia B J xa TL J 8 >< >: \u00f01\u00de In the above equations, xa, Kv, J, B, Kt and TL, respectively are the angular velocity, the armature velocity constant, the motor moment of inertia, the friction constant, the motor torque constant and the friction torque of the motor. In the proposed flapping bird, for each of the wings, a combination of 4-bar and 5-bar mechanisms is utilized as a power transmission system to independently create the flapping motion of each section of the wing, according to Fig. 1. The mechanisms are designed in such a way that the two links are attached to the crank (main gear) at two different radiuses of the crank and with a certain phase difference. The attached link to the crank at a smaller radius is connected to the first section of the wing, the attached link to the crank circumference is connected to the second section of the wing, and they produce the flapping motion of both wing sections with the crank rotation (Fig. 1). The source of flow supplied by the DC motor is considered as the input of the mechanisms. The kinematic equations of the flapping mechanism are extracted as follows. In order to extract the linear and angular velocities of the center of mass of link d1 connected to the first section of the wing, we have: V * A1 \u00bc d dt r1 cos k1 r1 sin k1 \u00bc r1 _k1 sin k1 r1 _k1 cos k1 \u00f02\u00de V * B1 \u00bc d dt e\u00fe l1 cos c1 h\u00fe l1 sin c1 \u00bc l1 _c1 sin c1 l1 _c1 cos c1 \u00f03\u00de In the above relations, V * A1 and V * B1 are respectively the linear velocity vectors of the points A1 and B1 (the beginning and end-points of link d1), r1 is the radius of the crank in which the d1 link is attached and l1 is the connection point of link d1 to the first section of the wing", " The 1-junction represents the constant flow passing through the armature, and the inductance and the resistance bonds are connected to it. The gyrator essentially acts as the energy converter from electrical to mechanical one, and is connected to the 1-junction which represents the angular velocity of the motor shaft. It should be noted that the core of the motor bond graph model is its GY, which represents the propulsive force generated by the electrical energy. In addition, the input voltage is modeled as the source of effort (Se) in the bond graph. The small gear mount on the motor shaft is connected to the crank (Fig. 1), and in fact, the source of flow is a velocity that comes from the DC motor, and is considered as an input to produce the angular motion of the crank. It is presented as a 1-junction with an I-element that expresses the rotational inertia of the crank around its own axis. The bond graph of the flapping mechanism of the wings is constructed of 1-junctions related to the rotational velocities of the cranks, the flapping velocities, and the linear and angular velocities of the connecting rods center of mass, which are connected to I-elements represented the mass and the moment of inertia of the components of the mechanism" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002914_978-3-030-60330-4-Figure7.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002914_978-3-030-60330-4-Figure7.6-1.png", "caption": "Fig. 7.6 Scheme of force interactions between the particles and kinematics", "texts": [ " In the initial state, they are concentrated in the lattice sites and the distance between the mass centers of the neighboring granules along the x-axis equals a (Fig. 7.5). As in the rectangular lattice studied in Chap. 3, when moving in the plane, each particle has three degrees of freedom: the displacement of the mass center of the particle with the number N = N(i) along the axes x and y (translational degrees of freedom ui and wi) and the rotation with respect to the mass center (rotational degree of freedom \u03d5i) (Fig. 7.6). The displacements of the grains are supposed to be small in comparison with the period a of the considered one-dimensional lattice. It is assumed that the particle N interacts only with the two nearest neighbors in the chain, themass centers of which are located at the distance a along the axis x from the particle N (Fig. 7.6). The central and non-central interactions of the neighboring granules are simulated, like in Sect. 2.1, by elastic springs of three types: central (with rigidity K0), non-central (with rigidity K1), and diagonal (K2). The points of junctions of the springs with the particles are in the apexes of the rectangle of the maximum area, ABCE, inscribed in the ellipse (Fig. 7.6). Each rectangle has the size h1 \u00d7h2, where h1 = d1/ \u221a 2 and h2 = d2/ \u221a 2. The elongations of the central springs are determined by the changes of the distances between the geometrical centers of the rectanglesABCE, and the tensions of other springs are characterized by the variations of the distances between the apexes of these rectangles. 154 7 Propagation and Interaction of Nonlinear Waves \u2026 Note that, in contrast to the previously considered rectangular lattice of ellipseshaped particles (see Chap" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000485_etfa.2019.8869079-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000485_etfa.2019.8869079-Figure8-1.png", "caption": "Fig. 8. A single layer and two closed tracks.", "texts": [ " It relieves the programmer from directly managing XML code and therefore avoiding human programming errors. The developer operates on a class level, while AutomationML Engine creates the correct CAEX schema file according to the CAEX schema [2]. The AutomationML Generator imports and stores data from the CSV file to the AutomationML file. Such data are later provided to a robot post-processor. The AutomationML Generator allows to upload points from the path and has the ability to group them into tracks and layers, Fig.8. A layer consists of a group of points which are at the same distance regarding the normal surface of manufacturing, while a track is a sub element of a layer representing the points associated to the deposition of material. The criteria to identify a track consists on analysing when occurs an interruption in material deposition, this is, when the extruder stops the deposition on a track in order to travel and start deposition on another one. Having the path separated into tracks and layers allows a better control on the manufacturing process" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002488_tmag.2020.2997759-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002488_tmag.2020.2997759-Figure1-1.png", "caption": "Fig. 1: Conventional Direct Metal Deposition Technique from [4]", "texts": [ " For instance, DED has a high material deposition rate and high material utilization. In addition, this technique is superior for repair and add-on based applications. This technique can also be used for the installation of wear-resistant materials on products for extending life [3]. The authors are with the Department of Mechanical and Mechatronics Engineering, University of Waterloo, Waterloo, Ontario, Canada, N2L 3G1. Corresponding Author: M.B. Khamesee, khamesee@uwaterloo.ca As it can be seen in Fig. 1, there is a very strong dependence on the use of a substrate for its successful implementation. The primary function of the substrate is to constrain the article being manufacturing. The part being manufacturing is fused to the substrate to ensure that there is no motion in the x-y plane. The part is subsequently built layer by layer in the z axis [5]. [5] also studies the build up of residual stresses during the manufacturing process. Due to the constraining feature of the substrate, these stresses cannot be relieved. Thus, the articles being built are vulnerable to fracture. The work shown in this paper aims to bypass the need for this substrate. This is done through the use of magnetic levitation principles to actively levitate the article being manufacturing and producing the necessary restoration forces to stably hold the object. Fig. 1 depicts a conventional system setup for this technique. It can be seen that the system is constrained to the use of a single nozzle for material deposition. However, by the incorporation of Magnetic Levitation principles, multiple nozzles can be employed to enable faster manufacturing capabilities. Use of eddy currents to generate forces has been a popular means of generating non-contact forces. [6] studies the use of eddy currents to generate stabilizing restoration forces on tumbling uncontrolled satellites, thereby showing the viability of the principle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001875_978-4-431-54595-8-Figure8.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001875_978-4-431-54595-8-Figure8.7-1.png", "caption": "Fig. 8.7 Model of a target object", "texts": [ " Next, let us introduce the condition of a target object that keeps the threepoint support consisting of two hands and a leg. We assume that the object does not fall down even if the robot leans on it with both hands. The imaginary ZMP (Vukovratovic et al. 2001) of the object must be in the support area of the object for it to move stably, but it is difficult to know the real ZMP of the object. Here, we assume a simple model that the object with the mass m is supported in a rectangular area, as shown in Fig. 8.7. Let FEz be the hand reflect force in the vertical direction. Assuming a moment equilibrium equation about OB, the stability of the object is defined as follows: mgLOB FEzLEB > 0 (8.6) where LOB denotes the length from the CoM position to the rear end of the support area and LEB is the length from a contact point of a hand to the rear end of the support 182 T. Arai and H. Kamide area. Given an equation of the moment about OF in the x-y plane, the pushing force FEx in the forward direction is expressed as follows: FExLEG mgLOF < 0 (8" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003249_tte.2021.3049466-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003249_tte.2021.3049466-Figure14-1.png", "caption": "Fig. 14. Distributions of the core loss. (a) CS-FTPM. (b) SC-RDPM. (c) DC-RDPM. (d) HS-RDPM.", "texts": [ " To sum up, with the employment of a dual-stator structure, the torque density is improved; meanwhile, the phaseisolation performance is impaired. However, the comparisons above show that the HS-RDPM is ahead of SC-RDPM and DC-RDPM in the phase-isolation performance. It indicates that the poor phase-isolation performance of RDPM can be improved by the proposed hybrid slots, proving the effectiveness of optimization in Section II-C. To compare the topologies further, the core loss distributions and PM loss distributions under rated conditions are simulated here, the results of which are shown in Fig. 14 and Fig. 15 respectively. Fig. 14 shows the distributions of core loss, while the specific values are displayed in Fig. 16. It can be concluded that: 1) the stator and rotor of SC-RDPM produce the maximum core loss in four topologies; 2) the DC-RDPM Authorized licensed use limited to: Dedan Kimathi University of Technology. Downloaded on May 18,2021 at 11:03:36 UTC from IEEE Xplore. Restrictions apply. takes the second place in core loss; and 3) the HS-RDPM has a similar value of core loss to CS-FTPM, both of which are minimum. In addition, Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001605_acs.langmuir.5b04012-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001605_acs.langmuir.5b04012-Figure1-1.png", "caption": "Figure 1. A spherical hydrogel held against a solid surface by capillary forces is illustrated. The symbol b is the radius of the contact line of the meniscus with the sphere, a is the radius of the circle of contact of the gel with the solid surface with which it is in contact, and h is the height of the line of contact of the meniscus with the sphere line above the solid surface with which the sphere is in contact.", "texts": [ " It then has its own supply of water, independent of the atmospheric conditions, over a time interval short compared to the time that it takes for the water to evaporate from the hydrogel into the atmosphere, even if the relative humidity of the atmosphere is lower than the value at which it is in equilibrium with the water in the hydrogel. While some of that liquid may flow away as the gel is pressed against a surface, some of it will remain attached to the edges of the gel forming a meniscus (illustrated in Figure 1). Here we are assuming that this occurs over a time short compared to the time required for the fluid contained in the meniscus and the hydrogel to evaporate, or the humidity is sufficiently large to suppress evaporation. This will result in an attractive Laplace pressure for a hydrophilic gel and surface, which will provide an adhesive force to pull the gel toward the surface. When the gel is pulled away from the surface by a force applied to the gel\u2019s top surface (the gel surface opposite the surface in contact with the hydrophilic solid), the water in the meniscus will form a capillary bridge13 between the gel and the hydrophilic surface, which opposes the attempt to pull the gel off of the surface", "2 will consider polyelectrolyte polymer hydrogels, in which because the polymers are charged, the swelling of the gels is dominated by counterion osmotic pressure inside the gels. Although it is well-known that the degree of ionization of monomers depends on the pH,20,21 effects of pH will not be studied here. Rather, we will assume that the solution already has the appropriate pH to ionize the polymers in the gel when we consider charged polymers gels. 1. Effects of Capillary Forces on a Neutral Hydrogel Hemisphere. When a hemispherical hydrogel swollen with water is pressed against a flat surface, as illustrated in Figure 1, water will be expelled. Some of the water may flow away, but some will remain to form a meniscus around the region of contact with the surface. If the surface against which the sphere is pressed is hydrophilic, the meniscus will provide an attractive capillary force which could hold the sphere pressed against the surface, even when we begin to pull the sphere off the surface. Let us assume that the adhesive contact term resulting from van der Waals forces between the sphere and surface is small compared to the capillary forces", " Then, the problem can be treated using the usual Hertz theory15\u221217 with the force pressing the sphere against the surface equal to the difference between the capillary force and the net repulsive contact force exerted on the sphere by the surface within the region of contact of radius a. The Laplace pressure within the meniscus is given to a good approximation by22 \u03b3 \u03b3= = \u0305P r h 2 L (1) where r is the smaller radius of curvature of the meniscus, h \u2248 2r\u27e8cos \u03b8\u27e9, where h is the height of the contact line of the fluid on the hemisphere above the solid surface (as illustrated in Figure 1) and \u03b3 \u0305 = \u03b3\u27e8cos \u03b8\u27e9, where \u03b3 is the surface tension of the expelled liquid and \u27e8cos \u03b8\u27e9 is an average value of the cosine of the contact angle of the gel and the surface. If a and b are small compared to the radius of the sphere R, where b is the radius of the contact line with the gel (see Figure 1), the capillary force on the sphere is given to a good approximation by \u03c0 \u03b3 \u03c0= = \u0305F P b h bc L 2 2 (2) Because hydrogels are porous, the Laplace pressure acts over the entire flattened surface of the sphere, in addition to the meniscus. Since when b \u226a R, h \u2248 (2R)\u22121(b2 \u2212 a2), a commonly used approximation15\u221218 for b,a \u226a R, Fc is given to a good approximation by \u03c0 \u03b3= \u0305 \u2212 F R b b a 2c 2 2 2 (3) In the Hertz model,15\u221218 the normal stress is assumed equal to = \u2212p r p r a( ) (1 / )0 2 2 1/2 (4) where p0 is the maximum repulsive pressure exerted by the surface on the sphere in the region of contact" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002857_s10015-020-00642-2-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002857_s10015-020-00642-2-Figure1-1.png", "caption": "Fig. 1 a The agent specification. b An example of the agent\u2019s 3D model overview", "texts": [ " This will lead to a significant improvement, albeit the step cannot be too large, because it will differ with the current policy too much. Future reward ( Gt ) is defined as below: (4)target = r + max a\ufffd Q(s\ufffd, a\ufffd; t). (5)Y DDQN t = r + Q ( s\ufffd, argmax a Q(s\ufffd, a; t); \u2212 t ) , (6)Gt = rt + rt+1 +\u22ef + k\u22121rt+k\u22121 = k\u22121 \u2211 i=0 irt+i. Computer simulations are conducted to utilize the algorithms and generate behavior. The agents and experiment environment are implemented in the physics engine Unity 3D. As illustrated in Fig.\u00a01a, each agent gets two motors, which can provide agents the locomotion ability and a certain thrust. As the sensory inputs for agents, we use a camera, eight distance sensors, and an electric compass. Camera input is a 128 \u00d7 128 pixel image with 90\u25e6 field of view located at the top of the agent. Eight infrared sensors can measure distance within 1.0 m are distributed on the agent. The two motors have 3 pattern outputs, in which the velocity of the two motors ( Vleft,Vright ) can be (2.0\u00a0m/s, 2.0\u00a0m/s), (2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002740_978-3-030-46886-6_5-Figure5.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002740_978-3-030-46886-6_5-Figure5.1-1.png", "caption": "Fig. 5.1 Schematic structure of the BFM", "texts": [ " The rotor broadly consists of three major parts: (1) the intracellular \u201cbasal body\u201d, formed by a set of linked ring structures (of \u223c50 nm maximum diameter) spanning the cell membranes and peptidoglycan (PG) layer, (2) the long (\u223c10\u00b5m) external flagellum, formed mainly by the protein FliC, and (3) a short (\u223c50 nm) flexible universal joint termed the \u201chook\u201d, which connects the basal body to the flagellum and allows the extracellular components to bend while rotating. The cytosolic part of the basal body, termed the C ring, is composed by multiple copies of the three proteins, FliG, FliM, and FliN, and plays an important role in torque generation and switching between clock-wise (CW) and counter clock-wise (CCW) rotation (Fig. 5.1). Rotation of the rotor, and thus the flagellum, is the result of the collective effort of a dozen unitary motors, the stator units, ion channels located in the inner membrane and firmly anchored to the rigid PG layer. Torque is produced at the interface between each stator unit and the section of the C ring facing the inner membrane and formed by FliG. As a microtubule is the track for kinesin, the FliG ring can be seen as a circular common track over which the different stator units independently step (Sowa and Berry 2008; Samuel and Berg 1996)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002479_physreve.101.052413-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002479_physreve.101.052413-Figure11-1.png", "caption": "FIG. 11. Schematic illustration of the influence of cargo levitation and torque diffusion on the proportion distribution of motors on different MTs. (a) Kinesins transport the cargo from MT1 toward the intersection. (b) Cargo levitation increases the proportion of kinesins on MT2 by changing the distances between the cargo and MTs. (c) A counterclockwise torque changes forces inside the kinesins and increases the proportion of kinesins on MT2.", "texts": [ " When N = 64, the average number of kinesins bind to MT1 is around 4.4. Since unbinding 4 kinesins simultaneously requires a large force, the cargo is likely to be trapped near MT1. Similarly, in the overpass situation, when an intermediate number of kinesins are attached to the cargo (N = 25), multiple kinesins bind to MT2 for a sufficient long time, which leads to the large passing probability (Fig. 10). Note that, the proportion distribution of kinesins on different MTs at the intersection depends on the cargo levitation and the external torque (Fig. 11). For example, when the cargo is levitated above, a larger proportion of kinesins are going to bind to MT2 because of the limited length of the cargo linkers. When a counter clockwise torque is applied to the cargo, the forces between kinesins and MT1 increase, which accelerates the unbinding from MT1 and modifies the proportion distribution of kinesins. Next, the transport dynamics on a bundle of two parallel MTs is studied. The distance between the surfaces of the two parallel MTs is set to be 15 nm in this study" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003867_s40516-021-00150-6-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003867_s40516-021-00150-6-Figure4-1.png", "caption": "Fig. 4 a Laser-arc hybrid weld bead geometry and b material flow in the gap", "texts": [ " Five welding parameters, namely, welding speed, wire feed speed, welding current, voltage, and laser power, varied along with the gap between the plates (0\u20131.2\u00a0mm), as shown in Table\u00a02. A total of 27 experiments was conducted. The experimental condition No. 4 was repeated three times to examine the process consistency. The weld bead was cross-sectioned and geometric attributes, namely, Reinforcement height (h), penetration at the top plate (d1), penetration at the bottom plate (d2), leg-length at the top plate (L2), leg-length at the bottom plate (L1), and throat thickness (t), as shown in Fig.\u00a0 4a, were obtained. The macrographs were analyzed for the material flow, as shown in Fig.\u00a04b. The tensile shear test was performed following JIS 2241 standard. Figure\u00a0 5a shows the sampled position of the tensile shear specimen. The fracture location was noted and analyzed for the mode of failure. Figure\u00a0 5b shows the typical failures in the weld metal (a few) and HAZ. Some of the samples also failed in the base metal. The joint strength, i.e., nominal stress, was calculated using the base plate cross-section area. 1 3 Ta bl e 2 R es ul ts o f l as er -a rc h yb rid w el di ng Ex pe rim en ta l c on di tio ns Re su lts N o M A G La se r G ap W el d ge om et ry fe at ur es (a ll di m en si on s i n m m ) Jo in t str en gt h (M pa ) W el di ng sp ee d (m /m in ) W ire fe ed sp ee d (m /m in ) C ur re nt ( A ) Vo lta ge ( V ) Po w er ( K W ) (m m ) H * d1 d2 L1 L2 t 1 1 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003296_j.measurement.2021.109056-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003296_j.measurement.2021.109056-Figure4-1.png", "caption": "Fig. 4. Experimental apparatus used to study the frequency beat problem when two rotors have very close rotation speeds.", "texts": [ " The angular displacement from the initial position of the test mass and the value of the correction mass for balancing can be determined considering a correction vector VC, (VC = XC@PC), such that, VC + V0 = 0. The correction mass value can be calculated by the product of the test mass with XC/X0. This cycle should be repeated until the appropriate level of balance is reached for the application. L.P. Ponci et al. Measurement 174 (2021) 109056 In order to study the beats frequency problem, we mounted two 1/4 HP power motors on the same base, both controlled by two frequency inverters, Fig. 4. Unbalanced discs were coupled to these motors in order to generate two distinct sources of vibrations, conceptually simulating practical applications such as the centrifugal separator and the dualrotor aeronautical gas turbine. Additionally, a NK820 Teknikao analyzer and a SDAV Teknikao proprietary software were used, in conjunction with a tachometer optical sensor with accuracy less than 0.1% and a piezoelectric accelerometer sensor with sensitivity of 100 mV/g which can be used at frequencies of 0", " Regarding the uncertainty of our experiments, we can state that the uncertainty of our experiments is below 2% according to Intermetro Calibration Laboratory from CGCRE in accordance with ABNT NBR ISO/IEC 17025. Initially, global measurements of vibration in mm/s (RMS) were performed. This speed unit has the characteristic of being better associated with the vibration energy that is given as a function of centrifugal force and mobility. Fig. 5 shows several measurements of the global levels of vibration in some situations of interest as a function of time. The piezoelectric accelerometer sensor is installed on the horizontal direction of the bearing of motor 2 near the unbalanced disk, according to Fig. 4. Thus, seeing Fig. 5, it is possible to clearly understand the beats frequency problem in vibration analysis and, consequently, in the dynamic balancing of mechanical systems like the ones addressed in this paper. When only one of the motors is running and the other one is off, whether motor 1 or motor 2 is running, it is possible to measure globally the levels of vibration as a function of time. When the two motors are running simultaneously close to each other in speed and sharing the same base, there is instability in the measurement of global vibration levels", " This allows performing a standard vibration analysis and dynamic balancing by separating the signals of interest from each source as a function of time, even considering the two motors connected simultaneously and knowing that one source of vibration generates interference in the other, which affects the global behavior. In Fig. 11 it is possible to see the signal as a function of frequency with the cursors defining the frequency bands to be filtered. In the proprietary software SDAV there is a digital frequency filter that allows separating the frequencies as can be seen in Figs. 12\u201314. The tachometer optical sensor with reflective tape and the piezoelectric accelerometer sensor were installed to work on the bearing and the disk of motor 2, see Fig. 4. Fig. 12 shows the filtered signal in the time domain from the delimitation of 2938\u20133030 rpm, which refers to the interaction of the effects of the vibration sources of the two unbalanced motors. It can be clearly observed the amplitude modulation of signal and the detection of reflexive tape by the tachometer optical sensor as detached by the markers shown in red. Fig. 13 shows the filtered signal in the time domain from the delimitation of 2989\u20133013 rpm, which refers to the source of vibration of motor 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure13.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure13.9-1.png", "caption": "Figure 13.9 Locking nuts (a) standard lock nut (b) slotted and castle nuts (c) \u2018Nyloc \u2019 nut (d) \u2018Aerotight \u2019 nut (e) \u2018Philidas \u2019 self-locking nut (f) torque lock nut", "texts": [ " The threads are designated by the diameter in inches followed by the thread series, e.g. 3 __ 8 BSW (Appendix 3). It is not usual to include the number of t.p.i. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 13 Joining methods 189 nut is screwed on, the arms are forced into their original position and the resistance of these arms gives a good grip on the thread, preventing it from working loose. A third type, known as a \u2018Philidas \u2019 self-locking nut, Fig.\u00a013.9(e), has a reduced diameter above the hexagon. Two slots are cut opposite each other in the reduced diameter and the metal above the slots is pushed down, which upsets the thread pitch. When screwed in position, the thread is gripped by the upset portion, preventing the nut from working loose. A fourth type, known as a torque lock nut, Fig.\u00a013.9(f), has the top part of the nut deformed to an elliptical shape which grips the thread as the nut is applied. This ensures close contact between the threads, preventing the nut from working loose. As well as the pre-applied adhesives already mentioned, threadlocking products can be applied separately to secure any threaded fastener against vibration and shock loads. Applied as easy flowing liquids which fill the gap between mating threads, they cure into a strong insoluble plastic. The joint formed is shock, leak, corrosion and vibration proof and depending on the grade used, can be undone using hand tools and by the application of\u00a0heat", "\u00a013.8(c). The liquid adhesive is encapsulated within the film. When the two threaded parts are assembled, the micro-capsules of adhesive are broken, releasing the adhesive, which hardens and provides a reliably sealed and locked thread. The simplest method of locking a nut in position is by applying a lock nut. Lock nuts are a little over half the thickness of a standard nut. When used in conjunction with a standard nut and tightened, the lock nut is pushed against the thread flanks and locked, Fig.\u00a013.9(a). Slotted and castle nuts are used in conjunction with wire or a split pin through a hole in the bolt to prevent the nut from working loose, Fig.\u00a013.9(b). Self-locking nuts are available which are easy to assemble and do not require a hole in the bolt or the use of a split pin. One type, known as a \u2018Nyloc \u2019 nut, Fig.\u00a013.9(c), incorporates a nylon insert round the inner top end of the nut. As the nut is screwed on, the nylon yields and forms a thread, creating high friction and resistance to loosening. deflected inwards and downwards. When the D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 13 Joining methods 13 190 bite into the shaft and cannot be removed without destroying it. This type is available up to 25 mm diameter and can be used on all types of material, including plastics" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001570_s12161-016-0420-y-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001570_s12161-016-0420-y-Figure5-1.png", "caption": "Fig. 5 a Cyclic voltammograms obtained for 0.5 mMVB-12 at the GCE (a), poly(PBHQ)/GCE (b), MWCNT/GCE (c), and poly(PBHQ)/MWCNT/GCE (d) in 0.1 M PBS at a scan rate of 50 mV s\u22121. b Cyclic voltammogram for a poly(PBHQ)/ MWCNT/GCE in 0.1 M PBS at pH 2.5.0 in the absence (a) and presence (b) of 0.5 mM VB-12, scan rate = 50 mV s\u22121", "texts": [ " As a consequence, scanning electron microscopy, Fourier transform infrared spectroscopy, and electrochemical impedance spectroscopy showed that the nanotubes were coated with a poly(PBHQ) layer. The rich electrochemistry of cyanocobalamin and related VB-12 derivatives has been extensively observed in a variety of media (Lexa and Sav\u00e9ant 1976; Lexa et al. 1980; Lexa and Sav\u00e9ant 1983). Electrochemical behaviors of cyanocobalamin (CN-CbA[Co(III)]) were analyzed at the bare GCE (Zheng and Lu 1997). Generally, electrochemical researchers indicated that VB-12 and its derivatives demonstrated rich redox chemistry which was concentrated on the cobalt atom. Figure 5a displays the CV patterns for different working electrodes at the 100 mV s\u22121 scan rate in phosphate-buffered solution of pH 2.5 containing 0.5 mM VB-12. At the unmodified GCE (Fig. 5, curve a), VB-12 exhibited two cathodic peaks at \u22120.569 V (P1) and \u22120.753 V (P2) which were assigned to processes of reduction Co(III) to Co(II) and Co(II) to Co(I) (Pala et al. 2014). As illustrated in Fig. 5 (curve b), at poly(PBHQ)/GCE, the cyclic voltammetric curve exhibits a single irreversible cathodic peak within the potential range of \u22121.0 to \u22120.5 V and VB-12 exhibits remarkably enhanced voltammetric response comparing with GCE electrode, with Epc = \u22120.65 V. As depicted schematically in Fig. 5a (curve c), a pair of extensive redox peaks, in which the \u0394Ep was 102 mV, was reached for 0.5 mM VB-12 on the MWCNTs/GCE. The anodic peak potential was located at \u2212634 mV, and the cathodic peak potential of VB-12 appeared at \u2212726 mV. At MWCNT-modified electrode, VB-12 displays remarkably enhanced response comparing with the unmodified GCE and the poly(PBHQ)/GCE. The MWCNTs/GCE increased the electrocatalytic activity area and promoted the electron transfer rate on the electrode surface. The same trial was done using the poly(PBHQ)/MWCNT-modified sensing electrode. The cycl ic vol tammetr ic response of poly(PBHQ)/MWCNTs/GCEs in 0.1 M PBS (pH 2.5) in the presence of VB12 (0.1 mM) includes a weak reduction peak at ca. \u22120.24 V (Fig. 5b), attributable to Co(III)/ Co(II) redox couple (forms VB-12 (or VB-12a) Bbase-on^ and VB-12r base-on, respectively) and a well-defined, intense reduction peak at \u22120.739 V due to the further reduction of Co(II) to Co(I), i.e., form B12s Bbase-off^ (Fig. 5a, curve d). These findings are in good agreement with other published results (Michopoulos et al. 2015). The oxidation peak at \u2212656 mV owes to the oxidation of Co(I) back to Co(II). The electrochemical reaction Co(III)/Co(II) occurs at potentials more positive than \u22120.3 V and cannot be seen in Fig. 5a. Finally, the electrode performance was compared with that of modified and unmodified electrodes. The sharp values of the anodic peak current of VB-12 that reached the poly(PBHQ)/MWCNTs/GCE (13 \u03bcA) is about 2 and 11 times higher than that at the MWCNTs/GCE (anodic peak current 6.17 \u03bcA) and unmodified GCE (anodic peak current 1.13 \u03bcA), respectively. The resolution between the peak potentials (anodic and cathodic peak) (\u0394Ep = 105 mV) approves the reversibility of the electrochemical reaction with n= 1 mol of electrons transferred in the electrode response", " The pH value of the voltammetric response of the sensing electrode toward 0.5 mM VB-12 was carefully investigated using CV. The effect of various supporting electrolytes such as phosphate buffer and BR buffer solutions on the VB-12 peak current was investigated using the poly(PBHQ)/MWCNTs/GCE electrode (voltammograms not shown). In this study, cyclic voltammograms for VB-12 were recorded between \u22120.3 and \u22121.2 V for each pH value. The relatively weak peak observed at \u22120.24 V corresponding to the reduction of Co(III) to Co(II) (see Fig. 5b) and the peak at \u22120.739 V versus Ag/AgCl belonged to the reduction of Co(II) to Co(I). The results showed similar redox behaviors in all the buffer types. However, the best-defined redox peak couples were observed in phosphate buffer with the highest signals at pH 2.5 (Fig. 7b). As can be seen in Fig. 7a, the response of the modified electrode was maximum at pH 2.5. The anodic and cathodic peak current gradually decreased from pH 2.5 to ~pH 8.0. Then, the peak current values remained almost constant at higher pH values up to pH=12" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002076_j.acme.2016.10.004-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002076_j.acme.2016.10.004-Figure4-1.png", "caption": "Fig. 4 \u2013 Structural diagram of the damping system mounted on the glass gatherer robot's lance: (1) flange mounting the lance to the cast iron block, (2) thrust plate, (3) outer spherical rings of the friction damper, (4) inner spherical ring of the friction damper, (5) spacer ring, (6) uniaxial tensometric dynamometer, (7) preloading back nut, (8) clamps, (9) clamping sleeve, (10) lock nut, (11) glass gatherer robot's lance [5].", "texts": [ " It is very important that the larger the actual contact area, the higher the coefficient of relative energy dissipation in structural connections. In the described solution, one of the outer rings was constructed in a way allowing for attaching of a dynamometer on the cylindrical outer surface thereof, whose task was to measure the preload force of the damper elements. After designing and constructing the friction damper, it was necessary to ensure the attachment of its elements to the lance in such a way as not to interfere in its structure. Fig. 4 shows a structural diagram of the friction damper mounted on the glass gatherer robot's lance. Damper spherical friction rings (3, 4) mounted between the fixing elements: thrust plate (2) and preloading back nut (7). The main task of these elements was to transfer the rotational movements of the lance cross-sections occurring during the transverse vibration of the ball gatherer' arm to the damper's rings. The entire damping system was attached to the lance using clamps (8). On the basis of research from literature [3,4,14] and research conducted at the Institute of Production Engineering and Automation of the Wroc\u0142aw University of Technology [10], it was found that the damping properties of the described structure may vary depending on the value of the preload force of the damper rings" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001008_3dp.2015.0010-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001008_3dp.2015.0010-Figure5-1.png", "caption": "FIG. 5. The alternative 3D object used for process model validation. Note: figure not to scale of the 3D object for better visibility. Color images available online at www .liebertpub.com/3dp", "texts": [ " The results show that the median values for warping, surface, and diameter dimensions were close to the predicted values, and deviations from the model prediction were 1%, 1.5%, and 1%, respectively. The deviation from the predicted printing time was highest, where the median varied by 3.8% from the prediction of the first optimization plot. After the validation of the optimization plots using the original button 3D object, the optimal parameters were used to print a geometrically different 3D object to investigate the applicability of the model predictions. Figure 5 shows the second model used to validate the results. This 3D object was selected because of the differences in shape and structure. Instead of a circle, the model had a linear inclined surface and sharp angles. In addition, the different wall thickness in the cubicle inside the model was one of the differentiating factors. The 3D object was 30 mm in length, 20 mm wide, and had a height of 6 mm. The results of the prints with the second 3D object show mixed results in terms of the accuracy to the model predictions", " One brief calculation example supports this assumption: the small, button-shaped 3D object has a total printed volume of 264 mm3, whereas the sharp-edged 3D object is 1568 mm3 in printed volume. Adding an artificial factor to the model equation for printing time determined in the experiments that calculates as the solid volume divided by the reference volume of the button, we find that the print time predicted by this linear extrapolation (1) already presents a pretty accurate estimation (compare the prediction of 1663 s vs. the actual 1710 s and 1707.8 s depicted in Figure 5). tprint \u00bc V solid 264 mm3 \u00b7 Model equation print time \u00bc 1568 mm3 264 mm3 \u00b7 Model equation print time \u00bc 1663 s (1) It remains to be analyzed how this prediction can be made even more accurate, for example, by implementing a factor for the complexity of the model and the cutouts that create traverse time of the nozzle where no printing is done. Further major results of the experiments are as follows: The printing speed hardly influences the overall time needed to finish a printed part. Even though the temperature of \u2013 260 C increases the degree of melt viscosity, thus improving layer extrusion, the heat can leave burn marks on the model" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001268_s0263574714002653-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001268_s0263574714002653-Figure4-1.png", "caption": "Fig. 4. (Colour online) Geometric model of the first PKL. g1 = [ G1Op ] Sb (4) l21 = [A1F1]S11 (5) l31 = [F1G1]S21 . (6)", "texts": [ " Link (4\u2032) has two revolute joints with perpendicular axes, with the platform (5) and link (4), respectively. In this section, the inverse kinematic model for each PKL is developed. This model will be used in order to compute the orientation error of the manipulator due to the passive joints clearances. Let Sb(Ob, x0, y0, z0) and Sp(Op, x0, y0, z0) two reference frames fixed, respectively, on the base and the platform. The origin Ob (Op) corresponds to the reference point of the base (platform), as shown in Fig. 4. The y0 axis is taken parallel to the line 1, z0 axis is pointing upward from the base to the platform, while x0 is taken according to the right hand rule. Let S11(A1, x11, y0, z11), the reference frame fixed on the arm (2). The z11 is parallel to the line A1F1 (F1 corresponds to the middle of the segment C1C\u2032 1), while x11 axis is taken according to the right hand rule, S31(F1, x31, y0, z31) a reference frame, which defines the orientation of the parallelogram (D1D\u2032 1C\u2032 1C1), x31 is parallel to the axis of the revolute joint connecting the forearms (3) and (3\u2032) to the body (4) and z31 is taken according to the right hand rule" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000320_j.jnnfm.2017.11.005-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000320_j.jnnfm.2017.11.005-Figure1-1.png", "caption": "Fig. 1. Bent-core molecule and star molecule.", "texts": [ " In [17], a three-level schema is proposed for the modeling of rodlike liquid crystals, applicable for both static and dynamic theory. Starting from the molecular model, one can derive the tensor model, then the vector model (Oseen\u2013Frank and Ericksen\u2013Leslie theory), with all the coefficients determined by molecular parameters and the energy dissipation retained. We have applied the approach to the static theory of rigid molecules with achiral twofold symmetry, two of which are bent-core molecules and star molecules (see Fig. 1). For molecular https://doi.org/10.1016/j.jnnfm.2017.11.005 Received 7 May 2017; Received in revised form 18 October 2017; Accepted 17 November 2017 \u204e Corresponding author. E-mail addresses: xu924@purdue.edu (J. Xu), pzhang@pku.edu.cn (P. Zhang). Journal of Non-Newtonian Fluid Mechanics 251 (2018) 43\u201355 Available online 21 November 2017 0377-0257/ \u00a9 2017 Elsevier B.V. All rights reserved. T interaction, we consider the Onsager theory, i.e. adopt the hardcore interaction that is determined by the molecular architecture", "29), we have = = \u2212 = + \u2212m m m m m m m m\u03b1 \u03b1 \u03b1 I I I I, , 1 ( ).1 2 3 2 1 3 3 11 22 22 1 2 11 2 1 (2.35) By (2.21) and (2.32), we deduce that = \u23a1 \u23a3\u23a2 + + + + + \u23a4 \u23a6\u23a5 m m m m m m m m m m m m m m m m \u03c4 c\u03b6\u03ba m I I I I I I : ( )( ) . vf 0 22 1 1 1 1 11 2 2 2 2 11 22 11 22 1 2 2 1 1 2 2 1 (2.36) We can see that the only difference in the above terms lies in the coefficients as functions of the moment of inertia, from which we can distinguish the bent-core molecules and star molecules. For a bent-core molecule (drawn in Fig. 1 left), the sphere centers are distributed uniformly and continuously on a two-segment broken line, where the length of each segment is l/2. Thus, \u0302\u03c1 is given by \u0302 \u0302\u0302 \u222b \u239c \u239f= \u239b \u239d \u2212 \u239b \u239d \u2212 \u239e \u23a0 \u2212 \u239e \u23a0\u2212 r r m m\u03c1 l s \u03b4 l s \u03b8 s \u03b8( ) 1 d 4 cos 2 sin 2 .l 2 1 2 l 2 (2.37) Substituting it into (2.15) and recalling (2.16), we obtain = =I l m \u03b8 I l m \u03b8 48 \u00b74 sin 2 , 48 \u00b7cos 2 .11 2 0 2 22 2 0 2 (2.38) For a star molecule (drawn in Fig. 1 right), the sphere centers also lie in a third line segment of the length l2. Thus, \u0302\u03c1 is given by \u0302 \u0302 \u0302 \u0302 \u222b \u222b = + \u23a1 \u23a3\u23a2 \u239b \u239d \u2212 \u2212 \u2212 \u239e \u23a0 + \u2212 \u23a4 \u23a6\u23a5 \u2212 r r m m r m \u03c1 l l s \u03b4 l s \u03b8 s \u03b8 s \u03b4 s ( ) 1 d ( 4 )cos 2 sin 2 d ( ) . l l 2 2 1 2 0 1 l 2 2 (2.39) Therefore, = = \u239b \u239d \u239c + + \u2212 \u239e \u23a0 \u239fI l m \u03b8 I m l l l l x 12 sin 2 , cos , \u03b8 C11 2 0 2 22 0 1 3 2 3 1 12 3 2 2 2 2 (2.40) where = \u2212 + x l l l l cos C \u03b81 2 2 2 1 4 2 2 2 is the m1-coordinate of the center of mass. For bent-core molecules and star molecules, the spatial diffusion matrix J derived from the Kirkwood theory (see Appendix) is diagonal in the frame (mi), \u2211= = J m m \u03c0D\u03b7 \u03b31 8 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure15-1.png", "caption": "Figure 15. Results of the meshing models: (a) Quadrilateral mesh (Nodes 90680, Elements, 31879); (b) Hexahedral mesh (Nodes 103098, Elements 36901); (c) Fine mesh (Nodes 160918, Elements 88625)", "texts": [ " The finite element model of the rotor is carried out with a mesh of 20351 elements for a total of 39208 nodes. The mesh of the disc and pad resulting from ANSYS software is presented in Figure 14. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. A test of convergence is designed to evaluate the influence of the mesh on the precision of the digital simulation. One tested four cases of mesh (fine, hexahedral and quadrilateral) whose characteristics are presented in Table 6. Figure 15 shows the meshing models of the couple disc-pads. According to Table 7, one notices that the equivalent maximum constraints of Von Mises increase, according to the number of elements of the grid. The maximum value of the equivalent constraint of Von Mises as well as the total deformation attacks corresponds d to the maximum of the elements of the grid are practically those which one meets in the literature. It is thus judicious to choose a refined grid, because the solution becomes more exact by increasing the number of nodes of the grid. a) Influence smoothness of the mesh For that, one considered a second type of mesh finer and refined in the friction tracks, (Fig.16) appears. The element used in this mesh is a SOLID 187 and the total time of the simulation is equal to 8331.328 (s). This new mesh (type M2) consists of 11 3367 elements TE with 4 nodes, that is to say 18 5901 nodes. It is thus much finer than the mesh M1, Fig. 15 (c) used up to that point. Figure 17 shows the various configurations of displacements of the order \u00ecm of the model according to time, while keeping the symmetrical form compared to the vertical median plane. The total deformation is reached at the end of braking and varies between 0 to 52,829 \u00ecm . On the model of the not-deformed inner pad one has a degradation of the colors ranging from yellow and green to the red when the critical value is located on the higher radial edge of the deformed pad, shown in figure 15 by red color.This is due to the elastic modulus of of pad which is lower Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. than that of the disc.For the disc, we find that the displacements are located only on the friction tracks and its external crown; they reach a maximum value equal to 19,108 \u00ecm at time t = 45 [s] that is to say 36% of the total deformation of the inner pad, On figure 18 one notices that the outer pad has the same behavior as the inner pad on the level of the contact zone, but its total deformation falls to 67,43%, that is to say equivalent to 35,62 \u00ecm " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003290_s10846-020-01277-y-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003290_s10846-020-01277-y-Figure1-1.png", "caption": "Fig. 1 Representation of the considered system composed of multiple robots physically interacting with a human operator; the reference frames \u03a3w , \u03a3h, \u03a3o and \u03a3r,i are reported", "texts": [ " In the same section, the formal statement of problem addressed in this work is provided. A centralized solution to this problem is detailed in Section 3 and is then extended to a distributed setting in Section 4. Finally, simulation results are provided in Section 5, while conclusions and future works are drawn in Section 6. 2Mathematical Background Let us consider a system composed ofN serial-chain mobile manipulators which tightly grasp a rigid object and a human operator that co-manipulates the same object as in Fig. 1. In the rest of the paper, when we refer to human interaction, we intend that the human operator exerts, through his/herhand, forces on the the object that is tightly co-manipulated by the multi-robot system. In this way, the human is able to modify the object motion according to his/her desired motion. In the figure, the following reference frames are defined: \u2013 \u03a3w is the world reference frame; \u2013 \u03a3o is the object reference frame; \u2013 \u03a3r,i is the reference frame attached to the end effector of the i th robot; \u2013 \u03a3h is the reference frame attached to the human arm end point" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002605_tia.2020.3009955-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002605_tia.2020.3009955-Figure8-1.png", "caption": "Fig. 8. (a) Voltage relationship of the current balance condition. (b) Voltage relationship of the current imbalance condition.", "texts": [ " When the imbalance occurs, the negative-sequence back-electromotive force (back-EMF) drawn in red is superimposed to the positivesequence back-EMF, leading to the generation of the negativesequence current, which brings the second-order torque ripples and additional loss. The current balance control is to inject V1_n \u2212 Vm_n into the phase voltages to offset the Eneg. Similar to the harmonic suppression, the current balance control can be added to the MI or FI side, which would have different effects on the capacitor voltage. Ignoring the very small VFp in Fig. 3, the comparison of Fig. 8(a) with (b) shows the impact of the negative sequence component. Compared with Fig. 8(a) in the current balance condition, the extra negative sequence voltage Vneg should be added to offset Eneg rotating at -2\u03c9 in the d1-q1 plane, as shown in Fig. 8(b). Two options to add Vneg to the inverter are drawn in blue (MI side) and red (FI side). Due to the controllability of the FC voltage, it is better to add Vneg to the FI side, which can reduce the output burden of MI without affecting the reactive power compensation capacity of FI. As the critical element of the FI, the selection of the capacitor should also be carefully considered. From the perspective of the cost and capacity density, the electrolytic capacitor is more suitable and almost all previous studies have chosen to use it" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002886_jae-209506-Figure19-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002886_jae-209506-Figure19-1.png", "caption": "Fig. 19. Distribution of magnitude of magnetic flux density for load state for 52 and 87 harmonics.", "texts": [], "surrounding_texts": [ "Calculation of losses in the core of the motor supplied from the inverter requires taking into account both the higher harmonics of the magnetic field generated in the motor by the basic harmonics of the supply n voltage and by the dominant higher harmonics produced by the inverter, whose contribution is particularly significant at low frequencies of the supply voltage. When calculating the efficiency of the motor using the analytical circuit model, calculations for all the dominant harmonics of the voltages generated by the inverter should be made and the sums of individual components of these losses should be taken into account. It should be emphasized that for each harmonic voltage of the inverter, additional losses in the stator core must also be taken into account due to the higher harmonics of the magnetic field generated in the motor. The work showed that for virtually all frequencies, even when supplying the motor with sinusoidal voltage, additional losses in the core caused by higher harmonics of the magnetic field in the un cor rec ted pro of ver sio n machine under load, even at low motor operating frequencies, are comparable to the basic losses (for the tested motor for a frequency of 10\u00a0Hz additional losses to basic losses is 1.24, and for 20\u00a0Hz - 0.9) while for high frequencies they are many times higher than basic losses (for 350\u00a0Hz the ratio of additional losses to basic losses is 7.68). The increase in core losses due to harmonics generated by the inverter depends on the proportion of these harmonics in the inverter voltage and is much higher for low frequencies (for the tested motor at 10\u00a0Hz the ratio of losses at inverter power to losses at power supply is 2.5, and for 20\u00a0Hz - about 1.3) than for higher frequency (for 350\u00a0Hz the ratio of loss at inverter power to losses at power supply is less than 0.1). In addition, when supplying with low frequency voltage, there are large magnetic flux density values exceeding the values for which it was possible to measure magnetizing characteristics and specific losses of the core material. This is a source of additional errors both for analytical and field-circuit simulation, despite the use of different extrapolation methods." ] }, { "image_filename": "designv11_22_0003861_s40430-021-03138-7-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003861_s40430-021-03138-7-Figure3-1.png", "caption": "Fig. 3 Bearing test setup Fig. 4 Bearing test illustration and sensor placement", "texts": [], "surrounding_texts": [ "For each set of bearing data, 16 time-domain features are extracted. Butterworth filter is applied to reduce the impact of noise and then features are normalized in the range of [0,1]." ] }, { "image_filename": "designv11_22_0000959_s00502-014-0272-3-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000959_s00502-014-0272-3-Figure17-1.png", "caption": "Fig. 17. Complete 3D FEM model of the AFIR-machine", "texts": [ " These loss simulations yield to the following magnet losses and temperature rises in the magnets for OP3: Massive magnets: Pd,M,OP3 = 527.9 W and \u03d1M,OP3 = 96 K Segmented magnets: Pd,M,OP3 = 70.7 W and \u03d1M,OP3 = 56 K Hence for high speed operation a segmentation of the magnets is necessary, as it was done in a similar way also for the RFM. The losses and therefore the temperature rise in the magnets are reduced drastically due to the segmented magnets. Three permanent magnet synchronous machines as radial flux machine with outer rotor, axial flux machine with internal stator and axial flux machine with internal rotor (Fig. 17) have been electromagnetically designed for the use as wheel hub drives for compact class electric vehicles. A special focus was given on the choice of the number of slots per pole and phase q in the stator. It appeared that a high harmonic leakage factor in the air-gap field due to the stator tooth coil winding can lead to higher efficiencies at high speeds, especially when a big field weakening area is required. The axial flux machines show slightly higher calculated efficiencies at high speed in comparison to the radial flux machine" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001851_s12206-016-0614-7-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001851_s12206-016-0614-7-Figure7-1.png", "caption": "Fig. 7. Schematic diagram of the ending contact in the modified model.", "texts": [ " However, the uncorrected model has defined the exact contact position but is unable to reliably represent the uncertainty of contact for the taper leaf spring of tandem suspension. Although the exact location of contact between adjacent leaves in the taper leaf spring of tandem suspension is not sure, the contact between adjacent leaves must have occurred in both ends of the taper leaf spring of tandem suspension. Therefore, in the modified model, the number of contact force used to represent the end contact is increased to four given the \"to point to surface\" thought to extend the scope of the end contact and make the end contact closest to the actual situation. Fig. 7 shows a schematic diagram of the ending contact in the modified model. Interleaf friction occurred when leaves are kept in contact with adjacent leaves at both ends of the taper leaf spring of tandem suspension and would result in the hysteretic characteristics of leaf spring. In the uncorrected model, friction force between two adjacent leaves was calculated by a simple STEP function. The transition between static friction and dynamic friction was not taken into account, which does not agree with the actual situation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003340_tia.2021.3058200-Figure22-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003340_tia.2021.3058200-Figure22-1.png", "caption": "Fig. 22 Equipotential and circumferential flux density distributions of 18/17- pole IFC-BFPMM at four typical rotor positions with only armature reaction excitation. (a) 0 elec. deg. (b) 90 elec. deg. (c) 180 elec. deg. (d) 270 elec. deg.", "texts": [ " However, their cogging torques are too small to calculate accurately by FE method and large lower cogging torque harmonics in the figure are due to numerical errors. Fig. 21 shows the UMF waveforms of both machines at the rated current. Although both machines suffer from UMF due to their odd rotor pole numbers, the UMF of the 18/13-pole IFCDSPMM is much smaller than that of the 18/17-pole IFCBFPMM. Moreover, the method of doubling the stator/rotorpole number combinations can be adopted in order to eliminate the UMFs of both machines. Fig. 22 and Fig. 23 show the equipotential and circumferential flux density distributions of both machines at four typical rotor positions with only armature reaction excitation. In the 18/17-pole IFC-BFPMM, since every two Tcore segments are embedded with a PM, the flux generated by the coil wound on the stator tooth inevitably goes through the PMs. By contrast, the stator of the 18/13-pole IFC-DSPMM is separated into six E-core segments, most of the flux produced by the coil wound on one tooth of the E-core segment can pass Authorized licensed use limited to: Rutgers University. Downloaded on May 19,2021 at 02:52:57 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2021 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. > REPLACE THIS LINE WITH YOUR PAPER IDENTIFICATION NUMBER (DOUBLE-CLICK HERE TO EDIT) < 9 through the other two teeth, which ensures that only little flux goes through the PMs. As shown in Fig. 22(d) and Fig. 23(a), when the rotor positions of the 18/17-pole IFC-BFPMM and 18/13-pole IFC-DSPMM are 270 elec. deg. and 0 elec. deg. respectively, their analyzed PMs have potentially the risk of demagnetization, because their directions of circumferential flux density are opposite to those of PMs. Further, the circumferential flux density distributions of the two machines generated by PM and 3 times of rated current excitation are shown in Fig. 24 and Fig. 25. The used magnet material is N35H. Because their working temperature may reach 120 \u2103, the demagnetization of the PMs will occur with the circumferential flux density lower than 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003610_s11665-021-05762-9-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003610_s11665-021-05762-9-Figure3-1.png", "caption": "Fig. 3 Schematic of rotary disk into the feeder unit", "texts": [ " Some other calcu- lations were conducted, such as dilution, powder catchment efficiency, and building volume. The laser power parameters (P) 350 [W], deposition speed (vf) 2000 [mm/min], and standoff distance (zf) 3,5 [mm] remained constant during all experiments. According to previous analytical and experimental studies, these values were selected as the most appropriate for Inconel 625 and nozzle characteristics. The powder flow delivered at the nozzle is controlled by the rotation speed of a disk containing a circular slot (rdisk, wdisk, h in Fig. 3). The rotation is measured in percentage of disk rotation speed (%RPM), related to mass flow [g/min]. Gas flows through the coaxial nozzle by three channels named Nozzle (N), Shield (S), and Carrier (C). Flow meters are installed in gas lines to assess that in [l/min]. The hopper\u2019s mass flow due to the gravitational force and introduced in a slot created a powder layer boundary that is transported with an angular velocity produced by the rotary disk. The mass flow for this system is described by Eq 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003021_tec.2020.3045883-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003021_tec.2020.3045883-Figure13-1.png", "caption": "Fig. 13. (a) The 8/6 SRM\u2019s geometry with motor\u2019s windings modeling. (b) Temperature distribution of one phase winding with 7 A phase current. (c) Temperature distribution of the winding with zero-torque current control when rotor pole is fully aligned with stator pole in phase A ( IA = 3A, IB = ID = 5A, IC = 7A). ( Tamb = 25oC)", "texts": [ " Temperature variation of the winding during 80 min of injecting 5A current in all phases, 15 min without current, 80 min with 7A current in all phases, 15 min without current are shown in Fig. 12. Predicted temperature by thermal network (proposed model) and reduced-model also, measured temperature are explained by TN, R, and M in figure legends, respectively. A 3D thermal FEM is given to proof the accuracy of the proposed thermal model for steady-state temperature of the winding. Maximum estimated error by using the 3D thermal FEM and the proposed model is 7% of the winding temperature. The temperature distribution of the windings for 7A phase current is illustrated in Fig. 13. In TABLE V, the mean relative and the maximum absolute errors of the calculated temperature are reported for each current by the proposed model and the reduced-order model in respect to the maximum winding temperature. A point of interest in electric motors is the motor insulation life. In this study, an SRM has been used as a traction machine in an EV when the battery charger, motor and drive are an integrated structure. Therefore, dynamic monitoring of the winding temperature is of great importance because of the motor insulation life" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure3.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure3.4-1.png", "caption": "Fig. 3.4 Pose hypotheses generation using a \u2018birthday attack\u2019-like approach: Random dipoles are inserted alternatingly into 4D relation tables. After a fiew processing cycles one can find dipoles with similar relations, which thus are geometrically congruent. Graphic taken from [4]", "texts": [ " . . . . . 11 Figure 3.1 Regular structure of a 3D mesh acquired by a laser line scanner. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Figure 3.2 Rotation and translation invariant features of a dipole . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 15 Figure 3.3 The assumption of a tangential contact between two oriented point pairs can be used to define a relative transformation ATB . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 16 Figure 3.4 Pose hypotheses generation using a \u2018birthday attack\u2019like approach. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 18 Figure 3.5 Scan of piston rods lying on a table (SICK LMS400) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 20 Figure 3.6 Visualization of the occurring problem when sheet metal parts are scanned from a single point of view . . . . . . . 21 Figure 3.7 Rotational and translational invariant features of a tripole . . . . . . . . . . . . . . . . . ", "4 is invariant w.r.t. rotation and translation. In [96, 97] a highly efficient method for generating likely pose hypotheses by exploiting the theory of birthday attack [95]\u2014an efficient cryptological strategy to generate two different documents with similar digital signatures (hash values)\u2014has been proposed. In the following, this approach is summed up. Random dipoles of A and B are chosen, and alternately stored in relation tables (i.e. hash tables), using the four features of a dipole as table indices (see Fig. 3.4). Under the assumption that the invariants are unique, on average only 1.2 \u00b7 n pairs have to be processed until a collision occurs. This will provide a run-time complexity of O(n) [97]. More precisely, the 4D relation tables (one per surface), and the four invariant features (Eq. 3.4) as table indices are used. This leads to the following search loop: 18 3 3D Point Cloud Based Pose Estimation 1. Randomly choose a dipole with a, c \u2208 A and calculate rel(a, c). 2. Insert the point pair into the model\u2019s relation table: RA[rel(a, c)] = (a, c)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002058_1.4035203-Figure12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002058_1.4035203-Figure12-1.png", "caption": "Fig. 12 Interface of the simplified system", "texts": [ "url=/data/journals/jvacek/935929/ on 02/09/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use respectively, to derive the associated natural frequencies and mode shapes. The two subsystems are represented in modal space and coupled according to the compatibility and equilibrium conditions at the interface. This reduced model is of high efficiency and especially can allow explicitly description of the velocitydependent friction. Interaction on the interface is described with a simplified model, as shown in Fig. 12. The bearing is represented by linear springs and viscous dampers, which are connected in parallel. The stern support substructure is illustrated in Fig. 13, where the beams are marked with nos. 1\u20134. The bearing housing is considered as a rigid body. 4.1 Compatibility Conditions and Boundary Conditions. Compatibility conditions include the continuity of displacements and the equilibrium of forces, which can be deduced according to Table 1 Parameters of the finite element model Shaft Length Radius L1 L2 6" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002058_1.4035203-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002058_1.4035203-Figure13-1.png", "caption": "Fig. 13 Support substructure", "texts": [ " The two subsystems are represented in modal space and coupled according to the compatibility and equilibrium conditions at the interface. This reduced model is of high efficiency and especially can allow explicitly description of the velocitydependent friction. Interaction on the interface is described with a simplified model, as shown in Fig. 12. The bearing is represented by linear springs and viscous dampers, which are connected in parallel. The stern support substructure is illustrated in Fig. 13, where the beams are marked with nos. 1\u20134. The bearing housing is considered as a rigid body. 4.1 Compatibility Conditions and Boundary Conditions. Compatibility conditions include the continuity of displacements and the equilibrium of forces, which can be deduced according to Table 1 Parameters of the finite element model Shaft Length Radius L1 L2 6.16 m 0.03 m 3.82 m 2.16 m Beam Length Width Thickness 0.33 m 0.07 m 2.36 mm Material Density Elastic modulus Poisson\u2019s ratio 8034 kg/m3 2.03 1011 Pa 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000482_systol.2019.8864752-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000482_systol.2019.8864752-Figure1-1.png", "caption": "Fig. 1: Quadrotor Configuration", "texts": [], "surrounding_texts": [ "I. INTRODUCTION\nThe recent years have witnessed many developments in the area of unmanned aerial vehicles fleet and formation control. This is due to the benefits of operating multiple UAVs to achieve tasks more efficiently compared to a single UAV in a wide range of applications such as investigating large regions for surveillance and rescue or large payloads transportation [1]. The main challenge in this field is to find the local interaction rules between these vehicles to achieve a global desired comportment. To deal with this issue, different formation strategies are proposed in the literature and summarized below.\nIn the leader-follower configuration, the desired trajectory is loaded in the leader followed by the other agents of the flight [2]. The main drawback of this configuration is that the entire formation depends on the leader, in addition to the non self-organization formation. Solutions to some of these drawbacks were proposed in [3] where an interactive leaderfollower configuration of multiple quadrotors is considered to solve the problem of one leader dependency. In the virtual configuration, each UAV in the formation follows its own trajectory, where trajectories are computed in a central controller and shared with the agents without interactions between them [4]. The main drawbacks of this approach reside in the computational and communication cost and the possibility of collision occurrence in the presence of perturbations. The behavior-based structure is based on the collective behavior of animals where each agent must follow some guidelines to accomplish the formation, including the collision avoidance, the velocity matching and the formation centering. Different works considered this structure in the\nThe authors are with the Lebanese University, Faculty of Engineering, Research Scientific Center in Engineering, Hadath, Lebanon. {cfrancis, hassan.shraim}ul.edu.lb\nlitterature [5], [6] and demonstrate its advantages in terms of self organization, scalability and distributed control of agents. On the other hand, centralized, decentralized and distributed control architectures can be distinguished, where the last one was proven to be the most successful architecture for autonomous flight formation control [7]. Each agent in this architecture has its onboard controller that uses its own information and that of its neighbors obtained from sensing devices on each agent or through communication.\nDifferent control architectures used for autonomous vehicle fleet control exist in the literature in addition to the theoretical tools used to synthesize the control laws. Some methods proposed the use of consensus protocol to achieve a defined fleet goal [8], where the theory of algebraic graphs is the main theoretical tool used to solve control problems. Other methods are based on optimization approaches where the controllers are designed by minimizing or maximizing an optimization criterion. The advantage with these approaches in the control of multi-agent systems is that the objective is formulated as an optimization problem, providing effective tools to find the optimal solution with respect to the considered criteria. A distributed control strategy of a fleet of quadrotors based on a leaderless strategy is proposed in [9]-[11] where the formation problematic is stated as an optimization problem solved by the particle swarm optimization (PSO) algorithm. The objective is to minimize a cost function for each vehicle that enables the whole system to converge to a target point in a predefined configuration while avoiding obstacle and collision between agents. According to the authors, the PSO has shown optimal performance with a minimum calculation time. However, as stated in [12], the PSO can converge untimely and be snared into a local minimum, which could be risky when dealing with systems such as UAVs. To overcome this drawback, other optimization algorithms could be used to minimize the cost functions.\nThe Artificial Bee Colony (ABC) is a population-based optimization technique capable of handling constrained optimization problems [13]. Its was compared to other known optimization methods and the results outmatched or matched those achieved using other techniques [14], [15]. Different works in the literature proposed the use of ABC approach to solve mobile path planning [16] and satellite formation keeping problems [17], and show the feasibility of applying this technique for path planning and optimization of feedback gain controller parameters with a good quality of the\n978-1-7281-0380-8/19/$31.00 \u00a92019 IEEE 80", "solutions and fast convergence. In this paper, we propose the use of the Artificial Bee Colony algorithm for the leaderless distributed control of a fleet of quadrotor UAVs. The optimization problem is formulated as in the work proposed previously in [9]. The results show that the ABC approach is efficient, fast converging and suitable for such applications.\nThe paper is organized as follows. Section II is dedicated for the problem formulation. The Artificial Bee Colony algorithm is presented in Section III with its application to the quadrotor UAVs formation control. The results are presented in Section IV. The paper ends with conclusions and perspectives in Section V.\nThe interaction topologies of m agents in a multi-agents systems can be described using the graph theory [18], where the system is represented by an undirected graph G = (V, E), V = {1, 2, ...,m} and E \u2286 {(i, j) : i, j \u2208 V, i 6= j} being the sets of nodes and edges respectively. In a quadrotor fleet, each node is equivalent to a quadrotor UAV and the edges represent the communication between each two quadrotors. Let A \u2208 Rmxm be the adjacency matrix with elements aij > 0 if (i, j) \u2208 E and aij = 0 otherwise.\nTo represent the quadrotors fleet in the Euclidean space, a position vector qi = [xi yi]\nT \u2208 R2 is associated to each node i in the graph. A set of neighbors of a vehicle i is defined by:\n\u039ei = {j \u2208 m : ||qi \u2212 qj || \u2264 l} (1)\nwhere l is the neighborhood scope. The required configuration of multiple UAVs in the fleet could be formulated as follows:\n||qi \u2212 qj || = ddij \u2200j \u2208 \u039ei (2)\nwhere ddij is the desired inter-distance between two UAVs i and j. In addition, collision between agents should be avoided. This constraint can be written as follows:\n||qi \u2212 qj || > c (3)\nwith c \u2208 R being the smallest acceptable distance between any two neighbors. In the matrix A(t), this constraint is formulated as:\n\u2200aij(t) \u2208 A(t) : aij(t) > c, with i 6= j (4)\nTo introduce this constraint in the cost function, a new function \u03b1ij(t) is defined between each quadrotor i and its neighbor j [9]:\n\u03b1ij(t) = 1 + e c\u2212aij(t) \u03c3 (5)\nThe quadrotor UAV is modeled as a rigid body, and its nonlinear model is given as follows:\n\u03b6\u0307 = v m\u03b6\u0308 = g + RU \u03b7\u0307 = W\u2126\nJ\u2126\u0307 = \u2126xJ\u2126 + \u03c4\n(6)\nwhere \u03b6 = [x y z]T and v are the center of mass position and velocity vectors in the inertial frame, \u03b7 = [\u03c6 \u03b8 \u03c8]T and \u2126 are the angles and angular velocities in the body-fixed frame. The quadrotor\u2019s mass is m and g = [0 0 \u2212 9.81]T . J is the moment of inertia matrix. W is the transformation velocity matrix and R is the rotation matrix from the body frame to the inertial frame. U is the thrust vector and \u03c4 is the torque vector given as:\n\u03c4 = (f2 \u2212 f4).l (f3 \u2212 f1).l\n\u03c41 + \u03c42 + \u03c43 + \u03c44\n (7)\nwith fi and \u03c4i being the propeller\u2019s thrust force and reaction torque respectively:\nfi = Kf\u03c9 2 i \u03c4i = (\u22121)i+1K\u03c4\u03c9 2 i\n(8)\nKf and K\u03c4 are the thrust and torque coefficients.\nThe Artificial Bee Colony optimization method was introduced in 2005 by Karaboga [19]. It is inspired by the searching behavior of honey bee swarm. The colony of artificial bees includes three types of bees: employed, onlookers, and scout. Employed agents arbitrarily look for food-source locations referred to us as solutions. Then, by dancing, they send the nectar amounts of this food-source, referred to us as the solutions qualities, to the onlookers staying in the dance area of the hive. The period of a dance is equivalent to the nectar content, fitness value, of the food-source currently being exploited by the employed bee. Hence, by watching various dances, onlooker", "bees choose the appropriate food source position according to the probability proportional to its quality. A good foodsource position catches most of the bees. If onlookers and scout bees find a new food-source position, they may become employed bees. Moreover, when the food-source position has been tested fully, the employed bee associated with it leaves it, and may change it status to an onlooker or a scout bee.\nIn the ABC algorithm, each search cycle includes three steps: \u2022 In the initialization step, the food is produced, the\nemployed agent is sent to it to measure the nectar amounts and calculate the fitness. It is modeled as:\nxij = xminj + rand(0, 1)(xmaxj \u2212 xminj ) (9)\ni = 1, ... , n with n being the number of food sources and j varies from 1 to the dimension of x vector. After updating the position of the food source as in (9), the fitness response is calculated as follows:\nFitnessi = 1\n1 + \u039bi (10)\nwhere \u039bi represents the value of the cost function at the solution i. \u2022 When the search step is done, food sources description is sent to the onlooker bees: nectar amounts (fitness) and position. An onlooker bee analyzes all the information collected and selects the best food location according to its nectar quantity probability calculated as:\npi = Fitnessi\u2211n i=1 Fitnessi\n(11)\n\u2022 If the food site is consumed, it will be forsaken and a new exploration is launched for a new food location. This is done using a set counter (limit).\nThe solutions generation and update process in ABC algorithm is shown in Fig. 2.\nThe mathematical formulation of the optimization problem for the UAVs formation application is presented below where the cost function for each UAVi is defined as in [9]:\n\u039bi(t) = \u03c1(||Pd \u2212 [qi(t) + h]|| \u2212 ddip) + \u2211q i=1 \u03b1ij(t)(||qj(t)\u2212 [qi(t) + h]|| \u2212 ddij)\n+ \u2211m k=1Obski(t)\n(12) with: i 6= j, q = card(\u039e(t)), \u03c1 >> 1, ddip is the desired distance between UAVi and rendezvous point. h \u2208 R2 is the desired reference point to be found at each step time for each UAV according to the rendezvous point Pd and the neighbors xj(t), with j \u2208 \u039e(t). The obstacle avoidance constraint is formulated by the term below introduced into the cost function:\nObski(t) = e c\u2212||qkobs(t)\u2212[qi(t)+h]|| \u03c3 (13)\nThe objective of the ABC algorithm in this problem is to compute the optimal vector h for each quadrotor i that minimizes \u039bi(t) such that:\nlim t\u2192\u221e \u039bi(t) = 0\u2192 lim t\u2192\u221e L(t) = Ld(t) (14)\nLd is the desired geometrical configuration around a rendezvous point.\nFinally, the next position at time t+ t0 for the agent i will then be:\nYi(t+ t0) = Y (t) + h (15)\nThe proposed UAVs fleet control method has been validated by simulations on Matlab under different scenarios, then the trajectory following is shown when the algorithm is tested on a quadrotor fleet simulator on Matlab. The parameters of the ABC used for simulations are presented in table I.\nFirstly, the ABC technique is applied on a single agent to generate its path from an initial to a desired position without and with an obstacle, where the location of the obstacle is considered to be known. The results of this scenario are shown in Fig. 3 and 4. Fig. 4 shows how the ABC" ] }, { "image_filename": "designv11_22_0003657_aero50100.2021.9438399-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003657_aero50100.2021.9438399-Figure4-1.png", "caption": "Figure 4: HOTDOCK symmetries, s identifies the rotation around the vertical axis.", "texts": [ " There are some other important considerations to take into account about the geometry of the HOTDOCK interfaces. First, they present a 90\u00b0 symmetry with respect to the perpendicular axis of coupling. Since each HOTDOCK presents the four petals on the upper face, there are four possible 3 Authorized licensed use limited to: California State University Fresno. Downloaded on July 01,2021 at 00:49:04 UTC from IEEE Xplore. Restrictions apply. configurations in which two HOTDOCK interfaces can be coupled, as shown in Fig. 4. While these symmetries are not relevant to the SMT-SMT connection, they influence the robot configuration required to grasp each SMT. Second, the geometric representation of a HOTDOCK is described by the following parameters: the diameter, the height, and the effective height, which accounts for the emerging portion that protrudes over the surface of the mirror tile. Third, HOTDOCKS can be of three different types, depending on their functionality. The simplest ones (\u201cmechanical\u201d variant) only provide structural support, others are capable of transmitting power and data to an SMT (\u201cpassive\u201d variant), and a third type additionally provides active fixation (\u201cactive\u201d variant)", " The approach with contact made by the robot using compliant control is represented with the function ContactController(x), where x is the goal pose. 4 Authorized licensed use limited to: California State University Fresno. Downloaded on July 01,2021 at 00:49:04 UTC from IEEE Xplore. Restrictions apply. Algorithm 1 Sequence Task Function Input: \u2022 Desired mirror tile smti to be placed in the assembly. \u2022 Connector coni j of the mirror tile smti by which the mirror tile will be grasped. \u2022 Symmetry s \u2208 0, 1, 2, 3 used for coupling the HOTDOCK (see Fig. 4). Output: \u2022 Complete sequence of pick, move and place for the mirror tile smti in the assembly ASM ASMNOT \u2190 (SMT \u2212ASM) 1: function SEQUENCETASK(smti, con i j , s) 2: assert(smti \u2286 ASMNOT and G = {\u2205}) 3: Pick(smti, con i j , s) 4: Place() 5: end function Algorithm 2 Pick and Place functions Given: G and ASM 1: function PICK(smti, con i j , s) 2: assert(G = {\u2205} and smti ASM) 3: xf \u2190GetPickApproach(smti, con i j , s) 4: Move(xhome, xf ) Performs Task1 5: ContactController(xconi j ) Performs Task2 6: Grasp() Performs Task3 7: G = {(smti, con i j)} 8: ContactController(xf ) Performs Task4 9: Recalculate TCP as CoM of smti 10: end function 11: function PLACE 12: assert(G = {\u2205}) 13: smtp, \u2190 G" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure3-1.png", "caption": "Figure 3. Full disc", "texts": [ " With this machining, the temperature of the bowl effectively decreases, but the heat gradients increase consequently in this zone. Those generate thermal stresses which explain the ruptures of bowl observed during severe experimental tests. There are two types of disc: full discs and ventilated discs. The full discs, of simple geometry and thus of simple manufacture, are generally placed on the rear axle of the car. They are composed quite simply of a full crown connected to a \u201cbowl\u201d which is fixed to the hub of the car (Fig.3). The ventilated discs, of more complex geometry, appeared more tardily. They are most of the time on the nose gear. However, they are increasingly at the rear and front cars, upscale. composed of two crowns - called flasks - separated by fins (Fig. 4), they cook better than the full discs thanks to ventilation between the fins, which, moreover, promote convective heat transfer by increasing the exchange surface discs. The ventilated disc comprises more matter than the full disc; its capacity for calorific absorption is thus better" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002110_iros.2016.7759680-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002110_iros.2016.7759680-Figure2-1.png", "caption": "Fig. 2. The proposed VR-based 3D mapping system.", "texts": [ " Through the user\u2019s guidance, the robot was able to explore the environment efficiently without failure. Although human-robot collaboration can improve the efficiency of the mapping task, there are still some limitations. For example, the interface between the user and the robot is not intuitive enough to provide good user experience and easy control. In this paper, we propose a novel framework to connect human and robot closely for 3D mapping through the virtual reality (VR) technique. The proposed human-robot VR-based 3D mapping system is shown in Fig. 2. An Oculus Rift goggle [16] is used to connect the human with the robot. The Oculus Rift has dual lenses to provide a stereoscopic 3D perspective and an IMU for head movement tracking. The mobile platform consists of a mobile robot base, a compact mini-computer and an RGB-D camera. The mobile robot base provides the ability to move around in the environment. The RGB-D camera is used to capture the color and depth images necessary for creating the 3D maps. To enlarge the field of view (FoV), the RGB-D camera is mounted on the pan-tilt units controlled by a control board" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002854_j.engfailanal.2020.105028-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002854_j.engfailanal.2020.105028-Figure5-1.png", "caption": "Fig. 5. Restraints of Connecting rod.", "texts": [ " Comparison of mesh before and after fillet treatment. Z. Pan and Y. Zhang Engineering Failure Analysis xxx (xxxx) xxx According to the actual working conditions, the piston pin is installed in a \u201chalf floating\u201d way. And the piston pin and the connecting rod are relatively fixed. To facilitate the analysis of the connecting rod assembly of which model is globally restrained at the big end crank pin. The wrist pin is restrained to only allow movement in the bore axis. And the restraining system allows free thermal expansion. As shown in Fig. 5. There are some interactions to consider except restraints. Complex interactions among the components in the model assembly have to be considered in the structural analysis [22]. Key considerations are as Table 4 follows: In the connecting rod assembly, the constraints between components are shown in Table 4. Bound constraints are used among other components except for the parts listed in Table 4. The load on the connecting rod is the main factor leading to the fatigue of the connecting rod. Therefore, the load needs to be obtained from the front engine numerical simulation model before the analysis" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001208_gt2014-26176-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001208_gt2014-26176-Figure11-1.png", "caption": "Figure 11 FEM model of spline joint", "texts": [ " From the analysis above, we can get the following conclusion. The non-linear characteristics of spline joint lateral stiffness is caused by the change of contact status of contact faces while the stiffness loss is mainly caused by the structural discontinuity, such as the slide and separation of contact surfaces. In order to prove the conclusion obtained from the stiffness model described above, a finite element model was established, simulating the stiffness change. The spline joint simulation model is shown in Figure 11. There are 64200 nodes and 54220 elements in the model. It has the same size and material of the test model described in following section. The property parameters of material are shown in Table 1. The solid element is SOLID185. The simulation model consists of two parts, which are connected by contact elements, CONTA174 and TARGE170. There are totally 4 contact surfaces, two centering surfaces and two axial pre-tightening surfaces. Utilizing different diameters, the model can simulate different contact tolerances, as shown in Figure 12" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002798_s00170-020-06152-6-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002798_s00170-020-06152-6-Figure4-1.png", "caption": "Fig. 4 Example of datum target point, datum target line, and datum target area specification", "texts": [ " According to the concept of \u201cRestricted DF,\u201d a DF can be defined as one located restricted feature or as one unlocated restricted feature, as illustrated in Fig. 3 a and b. When the DF is an area and its location is not specified, then the restricted DF indication shall be completed by an orientation plane, to remove the location from other feature (Fig. 3b). Moreover, by the concept of \u201cdatum target\u201d a specific portion is taken from a DF. As widely used practice in several industrial sectors, this \u201ctarget\u201d portion is nominally a point, a line segment or an area, totally located on the real workpiece (Fig. 4). Datum target along with \u201cmovable datum target\u201d designation is considered a powerful tool, particularly relevant to identified AM needs, that enables the designer to create datums from partial surfaces, from offset surfaces and as well as to create a complete datum system from irregular surfaces (Fig. 5). Finally, geometrically restricted datum designation can be also achieved by the application of \u201ccontacting features,\u201d a rather sophisticated concept included in ISO relevant standards, which is here recognized as particularly handy for AM needs and applications" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002013_cca.2016.7587978-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002013_cca.2016.7587978-Figure1-1.png", "caption": "Figure 1. A 6-DOF AUV with its reference frames", "texts": [ " In Section III the objectives of this paper are explained. Section IV discusses the proposed control, then followed by the stability analysis from Lyapunov candidate in Section V. The simulation results and the conclusion are presented in Section VI and VII, respectively. This paper considers \ud835\udca9 AUVs with notation \ud835\udca9 = 1,2, \u2026 , \ud835\udc41. The kinematic and the dynamic model of \ud835\udc56 vehicle is given in the following chapter. Kinematic model studies about the geometrical position of the AUVs relative to the body-fixed velocity. The illustration can be seen in Fig. 1 and the equation which is described by Jacobian matrix \ud835\udc3d(\ud835\udf02) is shown in the following form [16]. ?\u0307?i = [ \ud835\udc3d1(\ud835\udf022) 03x3 03x3 \ud835\udc3d2(\ud835\udf022) ] \ud835\udc63\ud835\udc56 with \ud835\udc56 \u2208 \ud835\udca9 (1) Vina Putranti and Zool H. Ismail are with Centre for Artificial Intelligence & Robotics, UniversitiTeknologi Malaysia, Jalan Sultan Yahya Petra, 54100, Kuala Lumpur, Malaysia. vinawep@gmail.com, zool@utm.my Toru Namerikawa is with School of Integrated Design Engineering, Graduate School of Science and Technology, Keio University, 3-14-1 Hiyoshi, Kohoku-ku, Yokohama, Kanagawa 223-8522, Japan" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002323_0142331219892115-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002323_0142331219892115-Figure1-1.png", "caption": "Figure 1. Structure and model of single robot system: (a) general structure of a manipulator and (b) two-link Euler-Lagrange model.", "texts": [ ", 2019) dx= f t, x, u\u00f0 \u00dedt+ g t, x, u\u00f0 \u00dedv, \u00f03\u00de where x is the state, and u 2 R is the control input, and v is an m dimensional standard Wiener process defined on a probability space fO,F ,Pg with O being a sample space; F being a s field, and P being the probability measure; fi : R 3 Rn 3 R! R and gi : R 3 Rn 3 R! Rm are the continuous functions and locally Lipschitz in the rest of the arguments. Definition 1. (Ji and Xi, 2006). For any given V (x) 2 C2 of (3), the differential operator L is defined as LV = \u2202V \u2202t + \u2202V \u2202x f + 1 2 Tr gT \u22022V \u2202x2 g , \u00f04\u00de where Tr denotes the matrix trace. Problem statement Consider the networked multi-robotic systems expressed by Euler-Lagrange model shown in Figure 1 on undirected communication topology in the presence of noise and time-delay. m1 and m2 stand for the masses of links 1 and 2; qi1 and qi2 stand for the the angles of rotation of links 1 and 2; l1 and l2 denote the lengths of links 1 and 2. The goal of this paper is to design a noise-tolerance control scheme such that the distributed tracking control consensus of the networked multi-robotic systems is achieved. Based on Figure 1, if the following equations are satisfied, that is, lim t!\u2018 k qi(t) q0(t) k = 0, 8 i= 1, 2, . . . ,N and lim t!\u2018 k _qi(t) _q0(t) k = 0, 8 i= 1, 2, . . . ,N , then the consensus of formation control can be obtained. Noise-tolerance control protocol design with communication delay In order to achieve the formation control consensus for multirobotic networks, a control protocol is designed for each follower robot as follows ui t\u00f0 \u00de= gMi qi\u00f0 \u00de _qi t\u00f0 \u00de+Ci qi, _qi\u00f0 \u00de _qi t\u00f0 \u00de Mi qi\u00f0 \u00de 3 ( Ari t\u00f0 \u00de+ X j=N i aij +sij _v t\u00f0 \u00de ri t e t\u00f0 \u00de\u00f0 \u00de\u00bd rj t e t\u00f0 \u00de\u00f0 \u00de + bi ri t\u00f0 \u00de r0 t\u00f0 \u00de\u00bd ) , \u00f05\u00de where i 2 1, 2, " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.16-1.png", "caption": "FIGURE 2.16 Braking tire.", "texts": [ " Hence, reducing the tread band material will result in lower resistance. The two tire geometrical parameters having an effect on rolling resistance are: \u2022 Tire radius \u2022 Aspect ratio (section height/tire width). Rolling resistance is decreased for a larger tire radius or a lower aspect ratio (low profile tires). Hence, smaller tires have a larger rolling resistance coefficient. However, such tires are usually used for lighter cars with a lower tire load and therefore lower rolling resistance force. Consider a tire under a brake torque, as indicated in Figure 2.16. The brake torque Mz must be balanced by moments due to a brake force Fx and the tire load Fz. The offset of the tire load in front of the wheel center increases with respect to the free rolling tire. The tire will experience a slip speed of the wheel with respect to ground, reducing the angular speed and therefore increasing the effective rolling radius Re. In the ultimate situation of a sliding nonrolling tire, this radius of rotation will become unbounded, with the center of the rotation moving to z5N. This means that, in general under braking conditions, the effective rolling radius Re,braking will exceed the unloaded radius. The total longitudinal shear stress in the contact area now consists of a part due to free rolling (dashed in Figure 2.16) and a superimposed shear stress caused by braking. As a result, the major part of the tire in the contact area is stretched due to the brake torque. Tread elements entering the contact area first try to adhere to the road surface, with the longitudinal deflection and therefore, the shear stress increasing linearly along the contact zone. At a certain point, the shear stress reaches the limits of friction (\u03bc \u03c3z with local road friction \u03bc and normal stress \u03c3z under Coulomb law) and the treads begin to slide. As a result, the shear stress drops down along the rear part of the contact zone. In a similar way as discussed for a free rolling tire, one arrives at a distribution of the peripheral velocity of treads (with respect to the wheel center), as shown in the bottom part of Figure 2.16. Note that sliding begins in the rear of the contact area and extends toward the front part of the contact area for increasing brake torque, until finally sliding is apparent along the full contact area. In case of a tire under driving conditions, the angular speed is increased and therefore, the effective rolling radius Re,driving is decreased. In the ultimate case of a spinning tire on the spot, the effective rolling radius has decreased to zero (no forward speed) and the point of rotation coincides with the wheel center" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.10-1.png", "caption": "Fig. 3.10 Sequences of two rotations. (a) Space-fixed, sz :: sy : 90\u25e6 rotation about the vertical axis sz, followed by a 90\u25e6 rotation about the horizontal axis sy. (b) Space-fixed, sy :: sz : 90\u25e6 rotation about the horizontal axis sy, followed by a 90\u25e6 rotation about the vertical axis sz. (c) Body-fixed, by :: bz : 90\u25e6 rotation about the body-fixed axis by, followed by a 90\u25e6 rotation about the bodyfixed axis bz. The final orientation is the same as in (a). Body-fixed axes and space-fixed axes are superposed because the size of the rotations in this example is exactly 90\u25e6", "texts": [ "9 may help to better understand the problem: how should we distinguish between a downward movement of the object by a rotation about the space-fixed axis sy (as shown in Fig. 3.9a) and a downward movement by a rotation about the rotated, body-fixed axis by (Fig. 3.9b)? 3Appendix A.3.3 contains the proof that the body-fixed representation of rotations uses the inverse (i.e., the transpose) rotation matrix compared to the space-fixed representation. 3.4 Combined Rotations 39 Mathematically, the difference between rotations in space-fixed coordinates and body-fixed coordinates lies in the sequence in which the rotations are executed. This is illustrated in Fig. 3.10. The upper column (Fig. 3.10a) shows a rotation of an object about sz by \u03b8 = 90\u25e6, followed by a rotation about the space-fixed axis sy by \u03c6 = 90\u25e6. Mathematically, this is described by bi = Ry(\u03c6) \u00b7 Rz(\u03b8) \u00b7 si (3.20) with \u03b8 = \u03c6 = 90\u25e6. Note: The rotation that is executed first is on the right-hand side, because this is the first matrix to act on the object to be rotated: Ry(\u03c6) \u00b7 (Rz(\u03b8) \u00b7 si) = (Ry(\u03c6) \u00b7 Rz(\u03b8)) \u00b7 si . (3.21) This leads to Rule 1: Subsequent rotations are written right-to-left. 40 3 Rotation Matrices Inverting the sequence of two rotations about space-fixed axes changes the final orientation of the object. This can be seen in Fig. 3.10b, where the sequence of rotations is inverted: the first rotation is about the space-fixed axis sy, and the second rotation about the space-fixed sz. This sequence is mathematically described by bi = Rz(\u03b8) \u00b7 Ry(\u03c6) \u00b7 si. (3.22) Equations (3.20) and (3.22) both describe rotations about space-fixed axes. However, they can also be re-interpreted as rotations about body-fixed axes in the reverse sequence: Eq. (3.20) can be re-interpreted as a rotation about the axis by by \u03c6, followed by a rotation about the body-fixed axis bz by \u03b8 (Fig. 3.10c). Figures3.10a and c demonstrate that rotations about space-fixed axes and rotations about object-fixed axes in the reverse sequence lead to the same final orientation. And Eq. (3.22) is equivalent to a rotation about bz by \u03b8 , followed by a rotation about the body-fixed axis by by \u03c6. A mathematical analysis of this problem can be found in (Altmann 1986). This can be summarized as Rule 2: A switch from a representation of subsequent rotations from space-fixed axes to body-fixed axes has to be accompanied by an inversion of the sequence of the rotation matrices" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001453_1350650114526582-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001453_1350650114526582-Figure1-1.png", "caption": "Figure 1. Schematics of the rolling bearing test rig used for this investigation. 1. Support bearing. 2. Test bearing. 3. Rack. 4. Free-moving bearing carrier.", "texts": [ " However, knowledge about the effect and mechanisms of additives on contaminated systems is missing. Therefore, this work was carried out to study the short-term effect of contamination on surface topography of raceways, vibration and acoustic emission signals with respect to the type of lubricant. The test samples were prepared in a laboratory environment and all the test setups were executed at room temperature. The investigation was carried out in an in-house-built rolling bearing test rig (Figures 1 and 2). The principle of the rig is based on a free-moving bearing carrier (pos. 4 in Figure 1) to be able to measure the sum of the friction torque from both test bearings (pos. 2 in Figure 1). The tested bearings are loaded with an axial load by a spring shown in Figure 2(a). Thereby the maximum adjustable axial load is 2.5 kN. The test rig is capable to run with a rotational speed between 100 r/min and 2500 r/min, due to a variable frequency drive. In the case of the 24 h tests friction torque and vibrations of the test bearings were measured. For the tests with duration of one week the rig was additionally equipped with acoustic emission sensors. The friction torque is measured by a force sensor connected to a torque arm (Figure 2(b))" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002979_0954406220979334-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002979_0954406220979334-Figure2-1.png", "caption": "Figure 2. Structure of the 2RT spherical parallel robot.31", "texts": [ " However, it requires the dynamic formulation of the robot alongside its kinematics and dynamics parameters to generate IDC terms. To not rely on the dynamic formulation and all kinematics and dynamics parameters, Partial Linearization IDC is recommended in this paper. Therefore, just the configuration dependent gravity term G of the robot is demanded to be compensated.43 Hence, it suffices to just apply a very slow trajectory with almost zero velocity and acceleration at any configuration in the workspace to the robot manipulator, such that the generated torque estimates the required gravitational term. Figure 2 depicts the structure of the 2RT spherical parallel robot in which, rotation of the active joints of the mechanism is about AO, and it is measured with respect to ZX plane about Z axis that is shown by h1 and h2 for each active joint, respectively. Furthermore, any point of the mechanism moves on the surface of a sphere for any input angles, and a spherical coordinate expresses its position. The orientation of the surgery instrument in spherical coordinates is considered through c and / in which c is rotation of OD about the X axis, while / is rotation of AD arc about Z axis and it is measured with respect to ZX plane" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002888_s12221-020-1016-0-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002888_s12221-020-1016-0-Figure1-1.png", "caption": "Figure 1. Experimental setup.", "texts": [ " The main thermal and mechanical properties of the biocomposite sheets are listed in Table 2. The glass transition temperature and melting temperature were obtained from differential scanning calorimetry (DSC) tests. The tensile strength, elongation at break and tensile modulus along the fiber yarn direction were determined by tensile tests according to ASTM D638. A three-axis CNC milling machine (VF-3SS, America, Hass) was used for SPIF of the composites. The fixture with a heating system is shown in Figure 1. The composite sheet was clamped by the blank-holder and a working area of 140 mm\u00d7140 mm was exposed. At the bottom of the clamped composite sheet, a backing plate was used to avoid the bending of the edge region of the composite sheet during the forming process. A forming tool with a hemispherical end and a diameter of 10 mm was controlled by the NC machining module. A constant tool speed of 3000 mm/min and zero spindle speed were assigned in the process. The hollow space inside the fixture was designed to store heat, which was insulated by glass cotton" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-FigureB.1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-FigureB.1-1.png", "caption": "Fig. B.1 Illustrates the transformation parameters of a pair of reference frames laid out according to Denavit-Hartenberg convention. (Illustration by Ethan Tira-Thompson, from WikiMedia Commons)", "texts": [ " In mechanical engineering, the \u201cDenavit-Hartenberg parameters\u201d (also called \u201cDH parameters\u201d) are the four parameters associated with a particular convention for attaching reference frames to the links of robot manipulators. In this convention, coordinate frames are attached to the joints between two links such that one transformation is associated with the joint, [Z], and the second is associated with the link [X]. The coordinate transformations along a serial robot consisting of n links form the kinematic equations of the robot, T = Z1 \u00b7 X1 \u00b7 Z2 \u00b7 X2 . . . \u00b7 Xn\u22121 \u00b7 Zn \u00b7 Xn, (B.1) where T is the transformation locating the end-link (see Fig. B.1). So going \u201cfrom the inside out\u201d, i.e. starting with the rotation-translation that does not affect any other joints: first the end-link is rotated about the x-axis, by an angle \u03b1. And it is translated along the x-axis by a distance r . (Note that the sequence in which this translation-rotation is executed has no consequence on the final position-orientation of the end-link.) This rotation-translation is described by the spatial transformation matrix Xi = \u23a1 \u23a2\u23a2\u23a3 1 0 0 ri 0 cos\u03b1i \u2212 sin\u03b1i 0 0 sin\u03b1i cos\u03b1i 0 0 0 0 1 \u23a4 \u23a5\u23a5\u23a6 , (B" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002528_s0263574720000429-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002528_s0263574720000429-Figure3-1.png", "caption": "Fig. 3. The soft robotic finger referring to human finger. (a) Prototype of the human finger. (b) Prototype of the proposed soft robotic finger. Referring to human finger, marker dots represent the joints and fingertip. Via marker dots, we can fit bending curve to represent finger state.", "texts": [ " In order to get rid of the noises, dilatation and corrosion transformation is performed on the binary image with 3 \u00d7 3 core twice. (3) Edge detection. Using the edge detection function of OpenCV, the boundary of each color region is obtained. (4) Image segmentation and scaling. The large image would increase the number of network parameters. Therefore, the boundary image requires segmentation and scaling from 640 \u00d7 480 to 224 \u00d7 115. Meanwhile, we use positions of joints and fingertip, referring to human finger (see Fig. 3(a)). Four marker dots are used to represent joints and fingertip (see Fig. 3(b)), which can fit bending curve of finger. The positions of marker dots are captured by outer camera as label image, which is processed in above steps as well. https://www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574720000429 Downloaded from https://www.cambridge.org/core. Uppsala Universitetsbibliotek, on 17 Jun 2020 at 08:46:41, subject to the Cambridge Core terms of use, available at After image processing, the processed images require to be normalized. For method A, the pixel values of scaled inner chamber image divide by 255", " Besides, the form of normalization of marker dots coordinates is as follows: (x, y) = ( x \u2212 xre lx , y \u2212 yre ly ) (1) where (xre, yre) is coordinate of reference marker dot. lx and L y are scale coefficient. Via normalization, convergence speed is faster and errors are decreased effectively. Then CNN19, 20 is applied for recognizing the inner chamber image. Because the features of the boundary information are obvious, so the simple CNN can easily recognize them. The structure of CNN is shown in Fig. 3(j). Each convolutional layer is followed by an activation layer, where rectified linear unit (ReLU)21, 22 is selected as activation function instead of Sigmoid22 and Tanh23 function. Since their gradient in the saturated region is close to 0, which is easy to cause the problem of vanishing gradient and reduce the convergence speed. On the contrary, ReLU\u2019s gradient is equal to 1 when value is more than 0, which is helpful to solve the convergence problem of network. In order to realize classification and regression, different classifiers and loss functions are used" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002730_cnm.3400-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002730_cnm.3400-Figure3-1.png", "caption": "Fig. 3. (a) Schematic diagram of the external fixator. (b) Equivalent model of the external fixator. The points Ai (i=1,2) and Bi", "texts": [ " (b) The distal ring moves 15 mm along the sagittal axis. (c) The distal ring moves 20 mm along the vertical axis. (d) The distal ring is rotated 20\u00b0 around the coronal axis. (e) The distal ring is rotated 15\u00b0 around the sagittal axis. (f) The distal ring is rotated 15\u00b0 around the vertical axis. The schematic diagram of the novel external fixator consists of two spherical joints, five prismatic (active) joints, six revolute joints, four universal joints, a proximal ring and a distal ring, as shown in Fig. 3 (a). To simplify the calculations involved in the fixator model, an equivalent kinematic model is established (a spherical joint and a revolute joint are equivalent to a composite spherical joint), as shown in Fig. 3 (b). The points Ai (i=1,2) This article is protected by copyright. All rights reserved. and Bi (i=1,2,...,4) are the centres of the spherical joints and universal joints, respectively. Revolute joint R3 is fixed on the proximal ring. The fixed coordinate system O-xyz is located at the intersection of the symmetry axis of the proximal ring and the axis of the revolute joint R3. The x-axis is along the axis of the revolute joint R3, and the y-axis is along the sagittal axis. The direction of the z-axis is determined according to the right-hand rule", " The moving coordinate system O'-x'y'z' is located at the intersection of axes of the three revolute joints R4, R5 and R6 in the constraint strut RPRRR. The x'-axis is along the axis of the revolute joint R6, the y'-axis is along the axis of the revolute joint R5, and the direction of the x'-axis and y'-axis are consistent with the x-axis and y-axis of the fixed coordinate system, respectively, in the natural position. The direction of the z'-axis is determined by the right-hand rule, as shown in Fig. 3 (a-b). (i=1,2,...,4) are the centres of the spherical joints and universal joints, respectively. The symbols Ri (i=1,2,...,6) represent the center of the revolute joints Ri (1,2,...,6). The length L1(L3) is the distance from the center of revolute joint R1(R2) to the point B1(B3). The length L2(L4) is the distance from the pointA1(A2) to B1(B4). The length L5 is the distance from the center of revolute joint R3 to the the center of revolute joint R4. The length li(i=1,2,3,4) is the distance from the point Ai (i=1,2) to the point Bi (i=1,2," ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure4.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure4.4-1.png", "caption": "Figure 4.4 \u2018Q-Max\u2019 sheet metal punch", "texts": [ " D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 Acknowledgements The author and publishers would like to thank the following organisations for their kind permission to reproduce photographs or illustrations. JSP Ltd. (Fig.\u00a01.7); Crown Copyright (Fig.\u00a01.8); Chubb Fire Ltd. (Fig.\u00a01.18); Desoutter Brothers Ltd. (Fig.\u00a02.22); Neill Tools Ltd. (Figs.\u00a03.10, 3.11, 3.14, 6.18, 6.23, 10.7, 10.8); Mitutoyo Ltd. (Figs.\u00a03.15, 5.33, 6.6, 6.9, 6.10, 6.13\u201317, 6.22, 6.26\u201328, 6.31\u201334, 6.36, 6.37, 6.39\u201341, 6.44, 6.45, 6.47); A.J. Morgan & Son (Lye) Ltd. (Figs.\u00a04.2, 4.8); Walton and Radcliffe (Sales) Ltd. (Fig 4.3); Q-Max (Electronics) (Fig.\u00a04.4); TI Coventry Gauge Ltd. (Figs.\u00a05.2, 5.5, 5.7, 5.16, 5.19, 5.21, 5.22); L.S. Starrett & Co./Webber Gage Division (Figs.\u00a05.11, 5.12); Verdict Gauge Sales (Fig.\u00a05.30); Rubert & Co. Ltd. (Fig.\u00a05.32); Thomas Mercer Ltd. (Figs.\u00a06.38, 6.42); Faro Technologies UK Ltd. (Fig.\u00a06.46); Draper Tools Ltd. (Figs.\u00a07.17, 7.19); Sandvik Coromant UK (Figs.\u00a07.22, 7.23); W.J. Meddings (Sales) Ltd. (Fig.\u00a08.1); Procter Machine Guards (www.machinesafety.co.uk) (Figs.\u00a08.3, 9.8, 9.9, 11.6); Clarkson-Osborne Tools Ltd (8", "9 22 0.7 24 0.6 Figure 4.1 Straight- and curved-blade snips D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 4 Sheet-metal operations 1 4 56 position while cutting takes place, and this also acts as a guard to prevent injury. These machines can be extremely dangerous if not used correctly, so take great care. When holes are to be cut in sheet metal, up to 16 SWG, this can be done simply and effectively using a \u2018Q-Max \u2019 sheet metal punch as shown in Fig. 4.4. A pilot hole is drilled in the correct position, the screw is inserted with the punch and die on either side of the sheet, and the screw is tightened. The metal is sheared giving a correct size and shape of hole in the required position. the linkage to the moving shear blade ensure adequate leverage to cut the thicker metals. Where larger sheets are required to be cut with straight edges, the guillotine is used. Sheet widths of 600 mm \u00d7 2 mm thick and up to 1200 mm \u00d7 1.6 mm thick can be accommodated in treadle-operated guillotines, Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000931_icep-iaac.2015.7111040-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000931_icep-iaac.2015.7111040-Figure1-1.png", "caption": "Fig. 1 Experimental apparatus", "texts": [ " In this paper, we focus on the surface roughness, contact pressure and hardness of materials from a number of factors. Then, the objective of this study is measurement of surface roughness and material hardness dependence of thermal contact resistance under low pressure condition. Experimental samples with several kinds of surface roughness (Ra = 0.2, 3.2, 12.5) and hardness are prepared, and thermal resistance between the samples is measured under several pressure conditions. II. EXPERIMENTAL APPARATUS Figure 1 shows experimental apparatus of thermal resistance measurement(4). The parts labeled No.1 and 2 in Fig. 1 are brass bars. Four thermocouples are located on each brass bar in equal intervals. The height of the brass bar is 45 mm. Thermal resistance of material is measured between these brass bars. The part labeled No.4 in Fig.1 is a cooling part. Water, which is kept at constant temperature, flows through this part. A film heater is located at part No.3 and heated by a power supply unit. The part labeled No.5 is polyimide block for insulating the heat flow from the heater to atmosphere. In following part, we call the brass bar at heated side as brass bar 2, the brass bar at cooled side as brass bar 1. The part labeled No.6 is measuring object. The details of the measuring objects will be shown in Chapter 3. The measurement technique of thermal resistance is below" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002414_pesgre45664.2020.9070682-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002414_pesgre45664.2020.9070682-Figure4-1.png", "caption": "Fig. 4: (a) A sample magnetic system; and its MEC (b) permeance network; (c) reluctance network", "texts": [ " Equation (29) can be written as \u2202g1 \u2202\u03a61 = R\u20321 + R\u20322 + R\u20324 Evaluation of \u2202g1 \u2202\u03c62 :- Partial derivative of (15) with respect to \u03a62 is \u2202g1 \u2202\u03a62 = \u2212R4 \u2212 \u2202R4 \u2202\u03a62 (\u03a62 \u2212 \u03a61) (30) Using the chain rule, \u2202R4 \u2202\u03a62 = \u2202R4 \u2202\u00b5r4 \u00b7 \u2202\u00b5r4 \u2202B4 \u00b7 \u2202B4 \u2202\u03a62 (31) The magnetic field density B4 at R4 is given by B4 = (\u03a61 \u2212 \u03a62) A4 \u21d2 \u2202B4 \u2202\u03a62 = \u2212 1 A4 substituting the value of \u2202R4 \u2202\u00b5r4 and \u2202B4 \u2202\u03a62 in (31) gives \u2202R4 \u2202\u03a62 = \u2212 R4 \u00b5r4 \u00b7 \u2202\u00b5r4 \u2202B4 \u00b7 \u2212 1 A4 Hence, \u2202g1 \u2202\u03a62 = \u2212R4 \u2212 R4 \u00b5r4 \u00b7 \u2202\u00b5r4 \u2202B4 \u00b7 1 A4 \u00b7 (\u03a62 \u2212 \u03a61) = \u2212R4 ( 1\u2212 B4 \u00b5r4 \u00b7 \u2202\u00b5r4 \u2202B4 ) = \u2212R4 \u00b5r4 \u00b5rd4 = \u2212R\u20324 Similarly, the remaining elements of Jacobian can be derived as \u2202g2 \u2202\u03a61 = \u2212R\u20324 \u2202g2 \u2202\u03a62 = R\u20323 + R\u20324 + R\u20325 The complete Jacobian matrix is[ G \u2032 l(\u03a6) ] = [ R\u20321 + R\u20322 + R\u20324 \u2212R\u20324 \u2212R\u20324 R\u20323 + R\u20324 + R\u20325 ] The elements of Jacobian matrix G \u2032 l resemble the corresponding element in the loop reluctance matrix R. III. CONVERGENCE BEHAVIOUR OF THE NEWTON-RAPHSON METHOD Convergence behaviour of the Newton-Raphson method while solving the non-linear MEC system is examined in this section. For simplicity, a sample magnetic system shown in Fig. 4 is considered with the following data. Core length lc = 0.17 m, air-gap length lg = 0.003 m, cross-section area of core Ac = 7 \u00d7 10\u22124 m2, stacking factor ks = 0.97, current in the coil I = 1 A, and number of turns in the coil is N = 600. The core is made of silicon steel material 35C250. The coil leakage flux and air-gap fringing flux are neglected in this study. Convergence behaviour of the Newton-Raphson method in both node and loop analysis is shown in Fig. 5, Fig. 6, Fig. 7, and Fig. 8. These figures are prepared using LATEX animation package to illustrate step by step iteration process. These animations can be played continuously or frame by frame using the control buttons provided below the figures. A. Node analysis The permeance network shown in Fig. 4b is solved using the Newton-Raphson method with two different initial conditions. In Fig. 5 and Fig. 6, the variation of residual function with the node \u00ac MMF is shown. At the true solution (F\u2217 = 596.39 A-t), the magnitude of residual function is zero. With the initial condition F(0) = 615 A-t, it is converging in three 978-1-7281-4251-7/20/$31.00 \u00a92020 IEEE 5 Authorized licensed use limited to: Auckland University of Technology. Downloaded on June 05,2020 at 18:32:56 UTC from IEEE Xplore. Restrictions apply", " The same permeance network is solved with the initial guess F(0) = 500 A-t, and the Newton-Raphson method shows diverging behaviour as illustrated in Fig. 6. In this case, the initial value and subsequently iterated value of the node MMF are located in the saturation region of residual function. From the above cases, it is clear that the initial condition is profoundly affecting the convergence behaviour of the Newton-Raphson method while solving the MEC system using node analysis. Hence, the initial value must be selected with the proper knowledge of the system; otherwise, it may not converge. B. Loop analysis The reluctance network shown in Fig. 4c is solved using the Newton-Raphson method with two different initial conditions. In both cases, the initial value of loop flux is located in the saturation region of residual function as shown in Fig. 7 and Fig. 8. The exact solution exists at \u03a6\u2217 = 0.1748 mWb. In loop analysis, shape of the residual function causes the NewtonRaphson method to converge even though the initial guess lies in the deep saturation region (i.e. B(0) = 2.14 T) as illustrated in Fig. 7 and Fig. 8. Irrespective of the initial guess value, the loop analysis shows stable converging behaviour" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure10-1.png", "caption": "Fig. 10. Failure mode and force-displacement curve for node in compression.", "texts": [], "surrounding_texts": [ "At present, very limited testing facilities that currently exist can respond to the needs of the daily increasing gridshell market, and a versatile and easily accessible testing equipment is in high demand. Although the designed test rig is customised for some specific load cases and geometry conditions (three-way node under tension, compression, shear and bending), the rules applied in the design process are general." ] }, { "image_filename": "designv11_22_0003469_s11665-021-05680-w-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003469_s11665-021-05680-w-Figure1-1.png", "caption": "Fig. 1 Schematic diagram of laser cladding process", "texts": [ " In the study, laser cladding was carried out with the help of a laser system (YLS-3000) and a magnetic field system (PEM5005) controlled with the parameters: the laser output wavelength of 1.07 lm, the laser scanning speed of 30 mm/s, the laser power of 1.50 kW, the laser spot diameter of 3 mm, the overlapping rate of 30%, the magnetic induction intensity of 1 T. High-purity argon with a speed of 3-4 L/min was used as shielding gas. The moving direction of laser cladding was parallel to the direction of the external magnetic field. Parallel laser tracks were fabricated to form a coating over the whole TC4 surface. The schematic diagram of the laser cladding process was shown in Fig. 1. After laser cladding, the phase compositions of HEA coatings were identified by Bruker D8 advanced x-ray diffraction at ranging from 20 to 90 , with Cu\u00c6Ka radiation, the tube voltage 40 kV, the tube current 30 mA, the scanning speed 5 / min. The coatings were characterized by CamScan MX2600FE scanning electron microscopy (SEM, Cambridge), with an acceleration voltage of 20 kV, equipped with an energydispersive x-ray spectroscopy (EDS, Oxford). Microhardness measurements of the HEA coatings were using by the HVST1000 Vickers hardness tester with a load of 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002150_mfi.2016.7849524-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002150_mfi.2016.7849524-Figure2-1.png", "caption": "Fig. 2. The test stand with modified control module (propulsion unit consisting of MN3110 BLDC motors with 10\u201d propellers)", "texts": [ " It consists of four coaxial propulsion drives presented in Fig. 1. Propulsion units can be divided into two types, i.e. CW-CCW and CCW-CW, with rotors rotating clockwise (CW) or counter-clockwise (CCW). 978-1-4673-9708-7/16/$31.00 \u00a92016 IEEE 418 In order to collect data for following analysis, experiments were performed on the test bench described in [15], with modified version of controller and software. Thanks to recent changes, the measurement circuit is more compact and reliable. All components of the mentioned device are showed in Fig. 2. Every propulsion unit consists of two MN3110 BLDC motors produced by RC Tiger Motors with 10\u201d carbon fiber propellers. The same components were used in Falcon V5 platform described in [1]. As can be seen, a steel arm does not correspond to the one used in real robot. This may cause a small discrepancies compared to the real vehicle and will be considered and analysed in ongoing research. This custom test bench allows to acquire measurements of generated thrust and power consumption (supply voltage and current) as a function of the PWM control signal parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001110_0020-7403(61)90036-4-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001110_0020-7403(61)90036-4-Figure1-1.png", "caption": "FIG. 1. Co-ordinates for Michell pad problem.", "texts": [ " I f p+ were the exact solution then equation (5) would automatically be satisfied. In the approximate procedure, equation (5) leads to n simultaneous equations for the a i, which are then solved. The \u00a2i are usually chosen to be simple polynomials or trigonometric series. This has the great advantage that the process consists mainly of manipulation of these simple functions. THE MICHELL PAD To assess the value of the method the classic problem first worked out by MichelP 6 in 1905 was attempted. The co-ordinate system given in Fig. 1 was used, the pad being square and the inlet film thickness twice that of the outlet. Written in non-dimensional form, the Reynolds equation becomes, with h = cx, ~ 2 p 2 v , 2~P . a ~ P x a ~ + 3 X ~ + A ~ y ~ + l = 0 (6) where P = pc21/6~?[g~l , x = I X , y = l Y . The boundary conditions are that P = O when { X = 1,2 +21 (7) A simple set of functions for P+ will be P+ = ( X - 1 ) ( X - 2 ) ( Y 2 - ~ ) ( a t + a 2 X + a 4 X 2 + a 6 y 2 + a ~ X a + a g X Y 2 + . . . ) (8) where all the cons tan t s in odd powers of Y have been pu t equal to zero because of the s y m m e t r y a b o u t the X-axis " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000150_j.wear.2018.12.037-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000150_j.wear.2018.12.037-Figure2-1.png", "caption": "Fig. 2. Experimental setup.", "texts": [ " The investigations were carried out on a Liebherr LC 120 hobbing machine with a modified tool spindle, which allows rotational speeds of n=10,000 rpm. This enables cutting velocities above the state of the art. The tool used in the experiments is a combination of a roughing hob and a fly cutter, which was presented by SCHALASTER for the first time [3]. With this experimental tool, the pre-machining and finishing operations in the hobbing process can be examined separately but in one setup. For this purpose, both a roughing hob and a fly cutter are mounted on the same tool shaft, compare Fig. 2. The pre-machined tooth gap is generated by the roughing hob, as can be seen on the left-hand part of the combination tool. After roughing, the tool is shifted to the finishing area. The finishing cut is realized with the fly-cutting process. During fly-cutting, the V-axis of the machine performs a continuous feed movement in order to realize all generating positions of the emulated hob. To assess the performance of the cutting materials, wear investigations were carried out with the presented fly-cutting trial" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001105_iros.2014.6943258-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001105_iros.2014.6943258-Figure3-1.png", "caption": "Fig. 3. Schematic of the stance leg controller. The controller consisted of a passive knee impedance (stiffness k and damping b) and hip torque \u03c4 . The hip torque was controlled by a constant velocity position trajectory command \u03c6c (Eq 1) with commanded hip stiffness and damping.", "texts": [ " Each leg was controlled using only local joint impedance commands and position trajectories, as in the implementation of the trot gait on the MIT Cheetah robot [16][17]. As the joint actuators were considered to be ideal within the torque and speed limits, joint stiffness and damping was commanded directly using proportional gains on the position and velocity feedback [19]; accurate limb impedance has been demonstrated on the MIT Cheetah [20]. This study focused on the impedance of the leg during stance. The stance leg control is shown schematically in Fig 3. During the stance phase, the knee impedance and leglength command were constant, thus the knee joint operated as a passive impedance [21]. The hip joint executed a constant velocity position trajectory described in Eq 1, where commanded hip position \u03c6c was a function of the stance time ts and commanded hip velocity \u03c6\u0307c. The stance clock ts was reset to zero at the beginning of each stance phase. \u03c6c(t) = \u03c6\u0307cts (1) The stance controller began with the knee fully extended and the hip at the commanded stance entrance angle with respect to the body" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003453_s12206-021-0334-5-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003453_s12206-021-0334-5-Figure6-1.png", "caption": "Fig. 6. Free body diagram of the proposed compliant mechanism.", "texts": [ " In PRBM, flexible links and the hinges are represented by torsional springs attached to the joints. If a mechanism is comprised of rigid links and flexures, deformations on the flexures are assumed to be small and linear in most cases. Dynamical model of the proposed mechanism is obtained using PRBM, vector closure loop equations, Euler\u2019s laws of motion and kinematic constrains. As in PRBM, flexures are considered to be torsional springs attached to the rigid links. Using the pseudo rigid body equivalent and free body diagram of the mechanism as depicted in Fig. 6, internal forces are defined as ( )( )1 1x xx ij i cii jF F m A+ +\u00b1 \u00b1 = (1) ( )( )1 1y yy ij i cii jF F m A+ +\u00b1 \u00b1 = (2) where i is the corresponding link, j is the link connected to the corresponding link (i) and xci A and yci A are the linear accel- eration of the center of corresponding link and m is the corresponding mass. Also sign of each force can be determined from Fig. 6. Moment equations of each link are expressed as 12 2 2 12 2 23 2 2 23 2 2 2 2 23 22*x y x yF r sQ F r cQ F r sQ F r cQ T T I \u03b8\u2212 \u2212 \u2212 \u2212 + + = (3) 23 3 3 23 3 3 34 3 3 34 3 3 23 34 3 3x y x yF r sQ F r cQ F r sQ F r cQ T T I \u03b8\u2212 + \u2212 + \u2212 \u2212 = (4) 34 4 4 34 4 4 45 4 4 45 4 4 34 45 4 4x y x yF r sQ F r cQ F r sQ F r cQ T T I \u03b8\u2212 + \u2212 + + \u2212 = (5) 56 5 5 56 5 5 45 5 5 45 5 5 56 45 5 5x y x yF r sQ F r cQ F r sQ F r cQ T T I \u03b8\u2212 \u2212 + + \u2212 + = (6) 67 6 6 67 6 6 56 6 6 56 6 6 67 56 6 6x y x yF r sQ F r cQ F r sQ F r cQ T T I \u03b8\u2212 \u2212 \u2212 \u2212 + + = (7) 7 7 7 7 7 7 67 7 7 67 7 7 7 67 7 7y x x yF r cQ F r sQ F r sQ F r cQ T T I \u03b8\u2212 + \u2212 + \u2212 \u2212 = (8) where ri is the half length of the ith link, is\u03b8 and ic\u03b8 are the sine and cosine of angle i\u03b8 measured from positive horizontal axis, Ii is the inertia, and \u00a8 \u03b8 is the angular acceleration" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001222_j.msea.2014.06.107-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001222_j.msea.2014.06.107-Figure1-1.png", "caption": "Fig. 1. Material parameters used for elastic linear kinematic plasticity assumption in (a) Warhadpande et al. [5] and (b) Bhargava et al. [6].", "texts": [ " However, not enough data were obtained for stress-controlled tests, and the stress\u2013strain responses used for calculations vary. For example, Warhadpande et al. simulated subsurface spalling during bearing operation [5] and Bhargava et al. calculated material deformation during rolling contact to predict the white-etching band morphology [6]. Although they both incorporated elastic linear kinematic plasticity, the material parameters were significantly different, especially for the yield limit (2.30 GPa versus 1.05 GPa) and the strain hardening slope (10 GPa versus 188 GPa) as shown in Fig. 1. According to tensile testing of 100Cr6 martensitic steels, the yield limit and strain hardening slope are within such ranges [7]. In this context, there is a need to obtain deformation data from stress controlled tests. In this study, a laboratory fatigue testing method is introduced to measure the plastic strain experienced by a martensitic 100Cr6 bearing steel. During the test, a repetitive compressive stress is applied to simulate the cyclic stress during bearing operation. The results would provide a database for the material parameters to be used in other calculations", " Now, (\u0394\u03f5C) is acquired by taking N to infinity in Eq. (5). Hence, \u0394\u03f5C \u00bc \u00f0\u0394\u03f5C;gage\u00deN \u00bc 1 \u00bc 0:225b3: \u00f06\u00de Finally, this is converted to its shear equivalent via \u0394\u03b3C \u00bcM\u0394\u03f5C : \u00f07\u00de For comparison, \u0394\u03b3C values from four literature sources were collected, these were obtained from calculations and experiments. Warhadpande et al. [5] employed the finite element method to estimate the cyclic shear stress\u2013strain hysteresis loop. They assumed that the material shows elastic linear kinematic plasticity with the mechanical parameters shown in Fig. 1a. \u0394\u03b3C was acquired from the resulting stabilised loop. Both Bhargava et al. [6] and Hahn et al. [12] employed the same material properties of normal stress\u2013strain hysteresis loop in Fig. 1b; \u0394\u03b3C was obtained from \u0394\u03b3C \u00bc ffiffiffi 3 p \u0394\u03f5C , according to Von Mises material assumption. Also, Hahn et al. [3] measured \u0394\u03f5C from cyclic torsion tests with a constant stress amplitude and converted to \u0394\u03b3C following the same expression. Christ et al. [4] employed a symmetric push\u2013pull fatigue test. The stress amplitude was taken to be equivalent to 1 2pmax and the corresponding p0 was calculated with Eq. (2) and \u0394\u03f5C was converted to \u0394\u03b3C via Eq. (7). In addition to \u0394\u03b3C , the maximum plastic strain (\u03f5p;max) has been acquired for each cycle during the testing" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.11-1.png", "caption": "Fig. 3.11 In gimbal systems, the axes of rotation are determined by the geometry of system. Both gimbals in this figure are in the reference orientation. Letbx,by,bz describe a body-fixed coordinate system. Left In a nautical (Fick) gimbal, the orientation of the object (the turn on the inner dial) is completely characterized by a rotation about the vertical axis bz by \u03b8N , followed by a rotation about the (rotated) horizontal axis by by \u03c6N , and a rotation about the (twice-rotated) dial-axis bx by\u03c8N .Right In a Helmholtz gimbalh, the orientation of the inner dial is characterized by a rotation first about the horizontal axis by by \u03c6H , followed by a rotation about the (rotated) bz axis by \u03b8H , and then a rotation about the dial-axis bx by \u03c8H", "texts": [ " If the rotated object is a camera or an eye, this roll rotation will not change the line of sight or gaze direction, but it will rotate the image. This sequence has first been used by the German doctor and physiologist Adolf Fick (Fick 1854), who worked on eye movements and who also invented the first contact lenses worn by patients. In eye movement research, the yaw, pitch, and roll angles for this sequence are therefore often referred to as \u201cFick angles\u201d. The yaw and pitch angles together determine the line of sight, and the corresponding direction is called the \u201cgaze direction\u201d. The left illustration in Fig. 3.11 shows a gimbal which corresponds to the nautical sequence of rotations. The angles of the nautical sequence will be denoted by the subscript \u201cN\u201d (\u03b8N , \u03c6N , \u03c8N ). The rotation matrix corresponding to the nautical sequence of rotations is Rnautical = Rz(\u03b8N ) \u00b7 Ry(\u03c6N ) \u00b7 Rx (\u03c8N ), (3.23) where the rotation matricesRx ,Ry,Rz describe per definition rotations about spacefixed axes. The discussion of Eqs. (3.20) and (3.22) defines the sequence in which nested rotations have to be written down: the first rotation (i", "24) This provides a convenientway to obtain the angles (\u03b8N , \u03c6N , \u03c8N ) from the rotation matrix R \u03c6N = \u2212 arcsin(Rzx ) \u03b8N = arcsin( Ryx cos\u03c6N ) \u03c8N = arcsin( Rzy cos\u03c6N ). (3.25) The nautical sequence is not the only sequence to describe the 3-D orientation of an object. Helmholtz (1867), another German physicist and physiologist from the nineteenth century, thought itwould be better to startwith a rotation about a horizontal axis. He characterized eye positions by a rotation about the horizontal interaural axis (i.e., the y-axis), followed by a rotation about the vertical axis, and then by a rotation about the line of sight, as shown in the right gimbal in Fig. 3.11: RHelm = Ry(\u03c6H ) \u00b7 Rz(\u03b8H ) \u00b7 Rx (\u03c8H ). (3.26) 3.4 Combined Rotations 43 The subscript \u201cH\u201d indicates that the angles refer to theHelmholtz sequence of rotations.One should keep inmind that the orientation of the object is characterized by the values of the rotation matrixR, andRnautical andRHelm only give different sequence parameterizations for the rotation matrix. But once constructed, the matrix is used in the same manner. Using Eqs. (3.14)\u2013(3.16) and matrix multiplication, we get RHelm = \u23a1 \u23a3 cos \u03b8H cos\u03c6H \u2212 sin \u03b8H cos\u03c6H cos\u03c8H + sin \u03c6H sin\u03c8H sin \u03b8H cos\u03c6H sin\u03c8H + sin \u03c6H cos\u03c8H sin \u03b8H cos \u03b8H cos\u03c8H \u2212 cos \u03b8H sin\u03c8H \u2212 cos \u03b8H sin \u03c6H sin \u03b8H sin \u03c6H cos\u03c8H + cos\u03c6H sin\u03c8H \u2212 sin \u03b8H sin \u03c6H sin\u03c8H + cos\u03c6H cos\u03c8H \u23a4 \u23a6 (3", " The projector can swivel on the disk left/right (\u03b8H ). And again, when \u03b8H = \u03c6H = 0 the projector is pointing straight ahead toward the screen. Task: What are the projector angles for the lower projector (\u03b8N , \u03c6N ), and for the upper projector (\u03b8H , \u03c6H ), if both should point at the target P = (hor/ver) on the screen? Solution: The rotation sequence for the lower projector corresponds to the two outer rotations of a nautical gimbal, and the sequence of the upper projector to the two outer rotations of a Helmholtz gimbal, respectively (see also Fig. 3.11). In both cases, the rotation about the target direction is unimportant, and \u03c8 in the equations for the nautical- and Helmholtz-rotation matrix can be set to zero. The direction to the target corresponds to the bx axis after the rotation, and the target point is the intersection of this axis with a plane parallel to the sy/sz-plane at a distance d. So for the lower projector, the target is at p = (d/ \u2212 hor/ver + below) b\u2032 x = p |p| . 50 3 Rotation Matrices (The sign before hor is negative, because the positive direction for \u201chorizontal\u201d on the screen is to the right, but the corresponding positive direction for \u201chorizontal\u201d for the projector is to the left", "11): the columns of the rotation matrix R are equivalent to the vectors of the body-fixed coordinate system [ bx by bz ] expressed in the space-fixed coordinate system [ sx sy sz ] . Thus, for eyemovementmeasurements with the search-coil method illustrated in Fig. 2.18, different values in the rotation matrix R indicate a different orientation of the eye-fixed coordinate system, i.e., a different orientation of the eye ball. Task: What is the orientation of an eye on a gimbal that is rotated 15\u25e6 to the left and 25\u25e6 down, if it is (i) a nautical gimbal or (ii) a Helmholtz gimbal? Solution: If an artificial eye ball on a nautical gimbal (Fig. 3.11a) is first turned 15\u25e6 to the left and then (about the rotated axis by) 25\u25e6 down, i.e., (\u03b8N , \u03c6N , \u03c8N ) = (15, 25, 0), its orientation will be given by the matrix Rnautical = \u23a1 \u23a3 0.88 \u22120.26 0.41 0.23 0.97 0.11 \u22120.42 0 0.91 \u23a4 \u23a6 . Putting an eye on a Helmholtz gimbal (Fig. 3.11b), and turning it first 25\u25e6 down and then 15\u25e6 to the left (about the rotated axis bz ), i.e., (\u03c6H , \u03b8H , \u03c8H ) = (25, 15, 0), leads to a different orientation of the eye: RHelm = \u23a1 \u23a3 0.88 \u22120.23 0.42 0.26 0.97 0 \u22120.41 0.11 0.91 \u23a4 \u23a6 . The orientation of the eye in the two examples is clearly different: on the nautical gimbal bz is given by (0.41, 0.11, 0.91), whereas on the Helmholtz gimbal it points in a different direction, (0.42, 0, 0.91). Interpretation: Experimentally, the three-dimensional orientation of the eye in space can be measured, for example, with induction coils (see Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002725_tmag.2020.3021644-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002725_tmag.2020.3021644-Figure10-1.png", "caption": "Fig. 10. Flux density distribution of the DFM-CMG corresponding to the prototype.", "texts": [ " And the combination of modulators in Fig. 9(c) is used as the output rotor. The 2D flux Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 20,2020 at 21:04:18 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. density distribution corresponding to the prototype is plotted in Fig. 10. It should be pointed out that the PM configuration, connecting bridges and mounting holes on OM may decrease the performances of DFM-CMG compared with the data in Table II. To verify that the DFM-CMG can improve the torque capability, CMGs I and II are also tested. For CMG II, only the IM is used as the output rotor. For CMG I, a back iron is glued to the inner side of PM ring, and only the OM is used as the output rotor. The experimental setup is shown in Fig. 11. The HSR of the DFM-CMG is driven by the servo motor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001279_s0263574714000149-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001279_s0263574714000149-Figure6-1.png", "caption": "Fig. 6. Natural coordinate\u2019s model.", "texts": [ " Each basic point or unit vector introduces three generalized coordinates called \u201cnatural coordinates\u201d. For example two basic points of a body introduce 6 generalized coordinates http://journals.cambridge.org Downloaded: 18 Mar 2015 IP address: 128.233.210.97 from which only 5 are independent due to the constant distance between two points of a rigid element; an element with two basic points and one unit vector introduces 9 natural coordinates, 6 of which are independent and so on. The natural coordinate\u2019s model of the same planar mechanism previously considered is presented in Fig. 6. All the elements are modeled using basic points: body 1, 2, 4 with 2 points each and body 3 and 5 with 3 points; points B, C, E and F are shared between the adjacent bodies. Therefore, for the 7 mobile basic points the number of the natural coordinates is 14. There are 9 constant distance constraints between the basic points and 4 constraints introduced by the closing loop joints D and GH, therefore the structural mobility can now be calculated as M = 2np,u \u2212 \u2211 dk \u2212 \u2211 ri = 2 \u00d7 7 \u2212 9 \u2212 4 = 1, (8) Accordingly, the kinematic model of the mechanism includes 9 constant distance equations and other 4 scalar equations for loop closure, in total 13 algebraic equations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure1.3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0001483_0954406215612814-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001483_0954406215612814-Figure1-1.png", "caption": "Figure 1. Generation of two conjugate profiles via the Camus theorem: a) starting position; b) conjugate-profile generation; and c) relative motion and validation.", "texts": [ "comDownloaded from This approach can be further extended to octoidal bevel gears by realizing that the return circle becomes a spherical curve of third degree and the line of action takes the form of a great circle of the fundamental sphere.20,21 The Camus theorem: Conjugate-profile generation Camus\u2019 theorem states that two conjugate profiles and s can be obtained as trajectories of a tracing point P that is attached to a suitable auxiliary centrode \", or as envelopes of a smooth curve that is attached to \", because the common normal to and s at their contact point P passes through the instant center of rotation P0, which is the tangent point between the centrodes l and l for the relative motion, as shown in Figure 1. This theorem can be applied to generate a pair of conjugate profiles from knowledge of the time-varying transmission ratio through the centrodes, which are traced by the instant center of rotation with respect to the two rigid bodies that roll with respect to each other by reproducing their relative motion. Thus, when the relative motion of a kinematic pair is assigned through the centrodes l and l, which are in contact at the instant center of rotation P0 and undergo a pure-rolling motion, the two conjugate profiles and s, which are in contact at point P, can be generated by applying Camus\u2019 theorem", " In fact, when a suitable auxiliary centrode \" is chosen as a continuous curve tangent to both centrodes at P0 and attached to another curve , both and s can be generated as envelope curves of during the purerolling motion of \" on l and \" on l, respectively. Therefore, the conjugate profiles and s are considered attached to the centrodes l and l, respectively, in order to allow the transmission of motion between the pair of rigid bodies according to the variable transmission ratio imposed by the centrodes, which roll with respect to each other, while the profiles and s remain in contact with sliding. In particular, referring to Figure 1(a), when a circle is chosen as auxiliary centrode \", a pair of conjugate profiles and s can be obtained as trajectories of a point P during the pure-rolling motion of \" on the centrodes l and l, respectively. In fact, from the starting position of Figure 1(a), in which \" is tangent to both centrodes l and l at the instant center of rotation P0, two different pure-rolling motions are considered, the first, of \" on l, to generate the profile , the second, of \" on l, to generate the profile s, as shown in Figure 1(b). The generic positions of \"0l and \"0l are shown in Figure 1(b), along with the positions P0l and P0l of the tracing point P. At the starting position, the generated profiles and s are tangent to each other at the contact point P; they remain constantly in contact during their relative motion, represented by the purerolling of l on l for a given pair of conjugate profiles and s, as shown in Figure 1(c). Camus\u2019 theorem can be also applied to the generation of the involute tooth profiles of circular and non-circular gears in the form of the well-known rack-cutter method. In fact, referring to Figure 2, when the auxiliary centrode \" becomes the tangent line t to both centrodes l and l, and its attached curve becomes also a straight line, a pair of involute conjugate profiles can be generated as envelope of , while their normal line N at their contact point P envelops the corresponding base curves, which are at UNIV CALIFORNIA SAN DIEGO on December 14, 2015pic" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure4.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure4.5-1.png", "caption": "Figure 4.5 Fly press", "texts": [ " These have a moving top blade, which is operated by a foot treadle, and a spring which returns the blade to the top of its stroke. The table is provided with guides, to maintain the cut edges square, and adjustable stops to provide a constant size over a number of components. When the treadle is operated, a clamp descends to hold the work in Figure 4.2 Hand-lever shears Figure 4.3 Treadle guillotine Where a number of components require the same size hole in the same position, it may be economical to manufacture a punch and die for the operation. The operation is carried out on a fly press, Fig. 4.5, with the punch, which is the size and shape of the hole required, fitted in the moving part of the press. The die, which contains a hole the same shape as the punch, but slightly larger to give clearance, is clamped to the table directly in line with the punch. When the handle of the fly press is rotated, the punch descends and a sheet of metal inserted between the punch and die will have a piece removed the same shape as the punch, Fig. 4.6. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 4 Sheet-metal operations 57 to allow for various thicknesses of material and can be made up in sections known as fingers to accommodate a previous fold" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.29-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.29-1.png", "caption": "FIGURE 2.29 Elliptic approximation of a tire friction envelope.", "texts": [ " In general, the drive force will be smaller than the brake force because it is bounded by the engine power. Furthermore, this enveloping curve will depend on speed on a wet road, especially in case of a significant amount of water (aquaplaning). A tire is not a rotational symmetric object, which explains the difference in size between \u03bcxp and \u03bcyp. This enveloping curve will be different for different tires. Let us assume the enveloping curve to be well approximated by an ellipse, as indicated by the outer curve in Figure 2.29. The right-hand part of this figure corresponds to driving (Fx. 0), whereas the left-hand side of this figure corresponds to braking. The outer ellipse describes the maximum shear force, which can be applied for a certain road friction and wheel load. The figure shows clearly (point A) that the side force in case of such a maximum shear force, under presence of a drive force Fx5\u03bcx Fz, will be less than \u03bcyp Fz. Likewise, applying a side force while braking or driving will reduce the longitudinal force, i", " We shall see later, when we plot Fx versus Fy for fixed slip angle, based on test data, that this approximation is rather rough. These elliptic approximations were used by Genta and Morello [10] to estimate the cornering stiffness under conditions of combined slip. In absence of a brake or drive force, the side force is indicated by Fy05\u03bcy0 Fz (i.e., in case of pure side slip). With maximum brake or drive force, the longitudinal friction coefficient is assumed equal to \u03bcxp. The internal ellipses in Figure 2.29 can therefore be described by the fol- lowing relationship between Fx and Fy: Fy \u03bcy0 Fz !2 1 Fx \u03bcxp Fz !2 5 1 \u00f02:66\u00de Assuming the slip angle to be small, such that the side force can be expressed as cornering stiffness times slip angle, the relationship (2.66) leads to C\u03b1;combined 5C\u03b1;pure ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 12 Fx \u03bcxp Fz !2vuut \u00f02:67\u00de with the C\u03b1,pure and C\u03b1,combined indicating the cornering stiffness in case of pure side slip and combined slip, respectively. We refer to Eq. (2.67) as the elliptic approximation of the cornering stiffness under combined slip. Following the approximation of the inner ellipse in Figure 2.29, we can conclude the following: \u2022 The side force Fy is a function of both \u03ba and \u03b1. For a fixed slip angle \u03b1, this side force has a maximum value at \u03ba5 0, which reduces with increasing j\u03baj (i.e., in case of either driving or braking). This also means that the peak value of Fy(\u03b1; \u03ba) versus \u03b1 decreases for increasing j\u03baj, as we observed previously. \u2022 Pure longitudinal slip characteristic behavior, as shown in Figure 2.17, shows a local peak value in Fx, followed by a decrease of Fx when \u03ba is further increased. This behavior is not shown in Figure 2.29, which indicates serious limitations in the elliptic approximation. Let us consider the polar diagram (Fx versus Fy for fixed slip angle), based on the tire parameters included in Appendix 6. The results are shown in Figure 2.30. Negative values for the slip angle correspond to negative Fyvalues. Indeed, one observes local maximum values for the longitudinal force Fx(\u03ba; \u03b1) versus \u03ba, with these values decreasing with increasing j\u03b1j (Figure 2.31). This behavior is similar to the lateral force for varying \u03ba, as observed previously", " Its value at vanishing slip angle tpa 3 ; \u03b1k0 is smaller than normally encountered (around a/2), see also Ref. [32]. We close this section with some remarks concerning the effect of brake/ drive force on the aligning torque and the approximations for combined slip contact force according to Eq. (2.71). Remarks 1. The polar plot of Fx versus Fy, as depicted in Figure 2.30 for the empirical Magic Formula, can also be derived for the physical brush model. Because the force characteristics based on the brush model saturate without decay for large slip, this polar plot will be similar to the one in Figure 2.29. We determined this polar plot for the brush model for Fz5 4000 [N] (see Figure 2.44). When we plot the aligning torque versus Fx, expression (2.82) leads to a plot that is symmetric in Fx, unlike Figure 2.33. This can be corrected by adding simple carcass flexibility to the brush model. This means that the entire carcass is pinned to the projected center of the wheel through springs acting in lateral and longitudinal direction with different stiffness values. Just as in the discussion on Figure 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000912_9781119016854.ch1-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000912_9781119016854.ch1-Figure8-1.png", "caption": "Figure 8. Additively manufactured Alloy 718 airfoil assemblies. This emerging technology can produce complex configurations directly from powder input stock.", "texts": [ " In addition to conventional subtractive processes to produce final component configurations, the emerging technology of additive manufacture has been applied extensively to Alloy 718. For these processes, Alloy 718 and variants are produced into powders that are recombined into desired form through localized melting and systematic and sequential build-up. Powder-bed and powder-feed processes have been successful in producing many unique configurations, often not possible by conventional processes or that would require high cost conventional processing. Alloy 718 has been extensively studied for additive manufacturing methods. [18, 19] Figure 8 shows a complex Alloy 718 airfoil assembly produced exclusively by additive manufacturing methods. Future Direction of Aerospace Material Requirements and Development Alloy 718 and its derivatives continued to have a bright future for emerging aerospace systems. The challenges largely continue to be the same, with component and system costs being extremely critical. Learning from the early development and evolution of Alloy 718 is important for any new material development that is aiming to become as ubiquitous as this significant alloy or as impactful as the lesser applied derivative alloys" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000538_0306419019887140-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000538_0306419019887140-Figure1-1.png", "caption": "Figure 1. CAD model of compliant parallel arm mechanism.", "texts": [ " Advances in additive manufacturing enable compliant mechanisms to be 3D printed using polylactic acid (PLA) or polyethylene terephthalate glycol (PETG, more flexible compared to PLA).37,38 Here, a compliant parallel arm mechanism is designed to demonstrate the displacement amplification and isolation of 2-DOF translational systems. The mech- anism consists of initially straight compliant beams, a primary and a secondary mass, a linear actuator, and two laser displacement sensors to record displacement of each mass separately as shown in Figure 1. The linear actuator is used to har- monically excite the primary mass with a frequency range from 1Hz to 25Hz. It is expected that the primary and secondary mass motion depend on the forcing frequency exerted by the actuator. The linear actuator is connected to the Arduino and controlled by a graphical user interface created in Matlab Simulink, enabling the user to change both amplitude and input frequency and record displacements in real time as presented in Figure 2. All parts of the mechanism and motor mount are 3D printed using a low-cost, home-type 3D printer (around $300) in PLA and assembled together", "003574 is fitted to the denormalized load\u2013deflection curve to represent the equivalent stiffness of the two-spring system attached to a slider KNL x\u00f0 \u00de \u00bc 0:008495x8 \u00fe 0:08284x7 \u00fe 0:3101x6 \u00fe 0:5649x5 \u00fe 0:5734x4 \u00fe 0:5437x3 \u00fe 8061x2 \u00fe 1:273x (8) Since all springs are identical and 3D printed using the same material, KNL1 and KNL2 will be equal. The load\u2013deflection curve has a linear relationship up to 2.5 cm of deflection due to the geometric constraints on both sides of the mechanism, as seen in Figure 1. Therefore, the flexible beams are not allowed to deflect more than 3 cm. As an alternative to the nonlinear springs, a linearized spring constant can be used by taking the slope of the curve between 0 and 2.5 cm. The linearized spring constant is calculated as Klinear \u00bc 0:4541 0:02513 \u00bc 18:07 ffi 18:0 N=m (9) Logarithmic decrement is an experimental way of finding the damping coefficient of an underdamped system.35 If the free response of an underdamped system can be acquired experimentally, then using the logarithmic decrement and damping ratio equations, equivalent damping of the system can be calculated" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002176_icra.2014.6907155-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002176_icra.2014.6907155-Figure9-1.png", "caption": "Fig. 9: (a) Pie chart for covered area by all triangles. (b) Histogram of covered area by each triangle.", "texts": [ " To evaluate a resulting triangulation quality or trace the trajectory of a navigation robot, we use AprilTag [16]. This measures the ground truth position, Pu = {xu, yu, \u03b8u}, of each robot u. The robots cannot measure or use the ground-truth position while executing our algorithms. Robots only know the two-hop local network geometry shown in Fig. 2b. Fig. 8 shows snapshots of triangulation. Over 8 trials using 9-16 robots, the average triangulated area is 1.5\u00b10.29m2. It takes 7.8\u00b12.1 robots to cover a unit area (1m2). The resulting triangulations are (\u03c1=3.6, \u03b1=0.36 rad)-fat. Fig. 9a shows that our triangulations cover about 91% of the region behind the frontier edges. The uncovered region is because the top-left and bottom-left corner in Fig. 8 are wall edges (incident on two wall robots), and are not expanded by navigating robots. Fig. 9b shows the distribution of area covered by individual triangles. The initial length of the base edge predicts the area of an ideal equilateral triangle should be 0.088m2, our triangles have a mean area of 0.13m2, with a std. dev. of 0.065m2. This discrepancy caused by the angle-based sensors; the robots cannot measure range, and therefore cannot control the area of the triangle they produce. We show this by studying individual triangle quality. Figs. 10a and 10b show our measurements of individual triangle quality: the distribution of minimum angle and maximum/minimum edge length ratio (MaxMin ratio) for each triangle" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002523_j.matpr.2020.05.322-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002523_j.matpr.2020.05.322-Figure10-1.png", "caption": "Fig. 10. Y axis Stati", "texts": [ " This reality brings the necessity of deciding the normal frequencies and mode states of the vehicle segments, and this applies unmistakably for the counter move bar. The mass and completed length of the bar are the remainder of the necessary yields from hostile to move bar examination. Clearly the mass of the bar ought to be limited, which is a general thought for all car segments. Likewise, the length is a significant parameter since it influences the mass and creation cost of the segment. Fig. 6 represents CATIA Model of an Anti-roll Bar and Fig. 7 shows Von Mises Stress of Anti Roll Bar. Fig. 8 shows Von Mises stress at a deformed Scale.Fig. 9.Fig. 10. Fig. 11. A power of 1000 N is applied toward the conclusion to make the connection as wound. Fig. (9\u201311) shows X Axis, Y axis and Z axis Static Nodal Stress. The stacking for the primary burden step, assurance of move firmness, is a known power, F, applied to the bar closes, in + Y heading toward one side and in \u2013 Y bearing at the opposite end. Figs. 12 and 13 shows the Stress intensity of the Bar and Von mises stresses of Bar. This redirection worth can be gotten by first plotting the \u2018\u2018DOF Solution - UY relocation\u201d form plot and afterward utilizing the question picker to peruse the uprooting estimation of the hub at the bar end" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000606_icems.2019.8922075-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000606_icems.2019.8922075-Figure1-1.png", "caption": "Fig. 1. Developed electromagnetic model of the high\u2013speed traction motor used for analysis.", "texts": [ " Performance characteristics in terms of torque\u2013speed characteristics and maximum energy density efficiency were compared for both machines. SPEED TRACTION PMSMS A 45 kW, 10,000 rpm high\u2013speed traction PMSM is considered for analysis. Two PMSMs with the structural and target performance parameters as defined in Table I were modelled using finite element based software incorporating different winding materials for analysis. The electromagnetic model thus obtained for the PMSMs is illustrated in Fig. 1 and the corresponding physical properties of copper and aluminum windings used are summarized in Table II [19]. 978-1-7281-3398-0/19/$31.00 \u00a92019 IEEE Both the aluminum and copper wound PMSMs developed have identical turns per slot, conductor diameter, and a 60% slot fill factor. Due to a 38% difference in material conductivity, the per\u2013phase resistance of the copper wound PMSM was found to be 0.044 \u03a9 and the aluminum wound PMSM was 0.068 \u03a9. However, replacing the copper windings with aluminum, resulted in a 12% reduction in weight and cost savings of almost 90% highlighting the inherent advantages of aluminum conductors, as illustrated in Table III" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002615_ieeeconf38699.2020.9389339-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002615_ieeeconf38699.2020.9389339-Figure3-1.png", "caption": "Fig. 3: The SeaBotix vLBV300 ROV, similar to the model used in this work, featuring six vectored thrusters, four in the horizontal plane and two in the vertical plane [15].", "texts": [], "surrounding_texts": [ "The simulated scenario involves a vehicle located at varying depths D within a water column of depth Dw = 50m, which attempts to remain stationary in the surge and heave directions, while subject to an oncoming wave train of varying significant height Hs. The system was modelled utilising fundamental concepts of wave theory [12] and dynamic equations for the vehicle itself [13] and subsequently the thrusters, to provide realistic behaviour with regards to response times and response to disturbances." ] }, { "image_filename": "designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure10-1.png", "caption": "Fig. 10. Screws. (a) Zero-pitch unit screw. (b) General screw.", "texts": [ " Appendix Motor representation of screws A screw, also called dual vector, can be represented as an ordered pair (a, ae) where the real part a and the dual part ae are three-dimensional vectors. By using the notation of VON MISES [34] a motor is a 6-vector a\u0302 \u2261 [ a ae ] (67) that contains the real and the dual part of a screw (dual vector). A zero-pitch unit screw (dual unit vector) describes an oriented line in space and is in motor notation given by u\u0302 \u2261 [ u ue ] = [ u r\u0303u ] , |u| = 1, r \u00d7 u \u2261 r\u0303u (68) with the unit vector of direction u and its moment ue with respect to the reference point O0, see Fig. 10a. The zero-pitch unit screw (68) fulfils the PL\u00dcCKER condition uT ue = 0. (69) A general screw comprises an oriented line in space with an associated pitch h (Fig. 10b), a\u0302 \u2261 [ a ae ] = [ a r\u0303a + ha ] = [ a ae\u22a5 + ae\u2016 ] , |a| = a. (70) Differential displacement of a screw In the following a differential displacement of a screw a\u0302 around the screw axis b\u0302 with the differential angle dv is considered, see also Fig. 11, with a\u0302 = [ a r\u0303aa + haa ] , b\u0302 = [ b r\u0303bb + hbb ] , |a| = |b| = 1. (71) The differential displacement of the screw a\u0302, da\u0302 = [ da d\u0303raa + r\u0303ada + hada ] \u2261 [ da dae ] , (72) is composed of the differential rotation of the vector a around the rotation axis b with the differential angle dv da = b\u0303adv (73) and the differential increment of the dual part dae of a\u0302" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002319_j.jsg.2020.104023-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002319_j.jsg.2020.104023-Figure1-1.png", "caption": "Fig. 1. Folding and boudinage under bulk plane strain. The competent layer, shown in red, is oriented perpendicular to one of the principal strain axes (X > Y > Z). Folds develop if the layer is oriented perpendicular to the X-axis. Boudins develop if the layer is oriented perpendicular to the Z-axis. Folds and boudins develop simultaneously if the layer is oriented perpendicular to the Yaxis (modified after Zulauf et al., 2003). Initial and final shape of samples are schematically depicted by the light grey cube and the dark grey cuboid, respectively.", "texts": [ "104023 Received 27 November 2019; Received in revised form 24 February 2020; Accepted 25 February 2020 Journal of Structural Geology 135 (2020) 104023 Given a rheologically stratified rock is deformed by dislocation (power-law) creep, the geometry of folds and/or boudins produced in the deformed layer(s) depends on various parameters (Hudleston and Treagus, 2010, and references therein): (1) thickness of layer(s), (2) shape and distribution of initial perturbations along the layer(s); (3) viscosity ratio between layer(s) and matrix; (4) stress exponent of the matrix and \u2013 more important \u2013 of the layer(s); (5) single or multilayer; (6) anisotropic viscosity of matrix and/or layer; (7) magnitude of bulk finite strain; (8) bulk strain rate; (9) bonding between matrix and layer (s); (10) bulk finite strain geometry (flattening, plane strain, constriction); (11) orientation of layer(s) with respect to the principal strain axes. Only few of these parameters can be accurately determined from naturally deformed rocks. The situation is relatively simple for cases where a single competent layer is oriented perpendicular to one of the principal strain axes (X > Y > Z) under bulk coaxial plane strain (Fig. 1). Single fold and single boudin axes will develop in the field of shortening and in the field of elongation, respectively. A large number of experimental and numerical models were produced under such conditions, where the layer is oriented either perpendicular to the long axis, X, or perpendicular to the short axis, Z, of the finite strain ellipsoid. This holds for classical plane-strain folds, with the fold axis parallel to the intermediate Y-axis (e.g. Biot, 1961; Biot et al., 1961; Ramberg, 1961; Hudleston, 1973; Cobbold, 1975; Abbassi and Mancketelow, 1992; Mancktelow, 1999; Schmalholz, 2006), and for plane-strain boudins, with the boudin axis (or neck line) parallel to the Y-axis (e", "5\ufffd, was almost half of the rotation of the corresponding passive plane. Under these conditions, the competent layer did not reach the field of elongation but remained close to the transition shortening/ reduced shortening (Fig. 11). The lower amount of shortening of the layer explains, why folding is less pronounced and folds are more symmetric. Folding of natural competent layers in incompetent matrix may result in thickened fold hinges and thinned limbs. These thinned limbs may be affected by boudinage (e.g. Fig. 20.8 in Ramsay and Huber, 1987, Fig. 1c in Druguet et al., 2009, Fig. 13.5 in Fossen, 2016). In the present studies, however, the finite strain of eZ \u00bc 70% was accommodated by 2 generations of almost homoaxial folds. This is the reason, why this high strain was not sufficient to initiate boudinage of the rotated fold limbs. Boudinage of fold limbs, however, is possible if shortening is accommodated by only one fold generation. Results of finite-element modelling, based on coaxial plane strain, suggest that under high bulk shortening and high viscosity ratio between layer and matrix, the least compressive stress (\u03c33) in the long limb of the fold becomes subparallel to the layer resulting in boudinage of the latter (Anthony and Wickham, 1978)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002379_0954406220917424-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002379_0954406220917424-Figure1-1.png", "caption": "Figure 1. Single-stage helical planetary gear set.", "texts": [ " Finally, both instead of the reliability and reliability sensitivity of the DTE fluctuations for the modification parameters were calculated with the limit state function. Computations for DTE of helical planetary gear set with tooth modifications In this section, the first stage gear train in a 750 kW wind turbine gearbox is used as an example to demonstrate the computation of the DTEs of gear pairs in a helical planetary gear set. A virtual prototype of the exemplary planetary gear set was developed in the ROMAX environment and is shown in Figure 1. As shown in Figure 1, the example system was composed of three central elements. For clarity, a sun gear is denoted as \u2018\u2018s,\u2019\u2019 a ring gear is denoted as \u2018\u2018r,\u2019\u2019 a planet carrier is denoted as \u2018\u2018c,\u2019\u2019 and a set of planet gears are denoted as \u2018\u2018pn\u2019\u2019 with n\u00bc 1, 2, 3. The carrier is set as the input component, and the sun gear is set as the output gear. The input rotation speed is set to be 22 r/min. Each planet gear is held to the rigid carrier through a rigid planet bearing, which is free to rotate relative to the carrier" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002900_j.tws.2020.107201-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002900_j.tws.2020.107201-Figure3-1.png", "caption": "Fig. 3. Stress distribution of single-corner tensioned membrane: (a) stress regions; (b) stress arc in region I; (c) stress arc in region II; (d) stress arc in region III.", "texts": [ " According to the previous solution of Timoshenko and Goodier [47] (Section 38, Chapter 4) for a point force on a wedge tip with unit thickness, the stress at a random point a relating to corner i (i = 1,2,3,4) in polar coordinates can be described as \u03c3r = k\u2217Pcos\u03b8 2ri , \u03c3\u03b8 = 0, \u03c4r\u03b8 = 0 (1) where \u03b8 is the angle between the line from corner i to point a and the diagonal line on which corner i exists, ri is the distance from corner i to point a, \u03c3r is the radial normal stress component, \u03c3\u03b8 is the circumferential normal stress component, \u03c4r\u03b8 is the shear stress component, and k\u2217 is the constant coefficient for \u03c3r. Eq. (1) indicates that the points which have the same value of \u03c3r form a stress arc in the membrane. In Fig. 3(a), the diameter of a stress arc is denoted as 2r0, and the distance from corner 1 to point a is denoted as r1. Therefore, r1 can be expressed with 2r0cos\u03b8, followed by being introduced into Eq. (1), which furtherly indicates \u03c3r has a constant value of k\u2217P/4r0 on the same stress arc. In the following, the stress distribution of the membrane with single tension load at corner 1 is illustrated. The stress arc is a part of a circle which crosses point a and corner 1, and its center which is denoted as O0 is on the diagonal from corner 1 to corner 3. The radius of the stress arc is denoted as r0. The angle between the direction of \u03c3r and the positive xaxis is denoted as \u03bb. According to the continuity of stress arc, the membrane is divided into three regions. In region I, a stress arc is continuous as the arc 5-6-7 in Fig. 3(b). In region II, a stress arc is divided into two continuous arcs in the membrane as the arc 5-6-9 and the arc 7\u20138 in Fig. 3(c). In region III, a stress arc is divided into three continuous arcs as the arc 5\u201311, the arc 9-6-10, and the arc 7\u20138, which are shown in Fig. 3(d). In a square membrane, region II does not exist B. Li et al. Thin-Walled Structures xxx (xxxx) xxx because the arc 5-6-9 in Fig. 3(c) will be divided into two arcs, thus being included by region III. As shown in Fig. 3(b), the distance from corner 1 to point 5 is denoted as l1, and that from corner 1 to point 7 is denoted as l2. The angle between line 1\u20136 and line 1\u20135 is \u03b11, and that between line 1\u20136 and line 1\u20137 is \u03b12. For the arc 6\u20137 in region II in Fig. 3(c), it includes three arc sections: arc 7\u20138, arc 6\u20139, and arc 8\u20139. The inscribed angles of arc 6\u20139 and arc 8\u20139 are denoted as \u03b71 and \u03b72, respectively. In Fig. 3(c), the middle point of the line segment 8\u20139 is denoted as M1, the distance from point M1 to point 9 is denoted as h1, and the distance from point 9 to the diagonal 1\u20133 is denoted as d9. Similarly, the location of M2 and the parameters h2, d10, \u03b73 and \u03b74 are defined and labeled in Fig. 3(d). No matter in which region a stress arc is located, the stress resultant on the same stress arc within the membrane is in equilibrium with the external tension load at corner 1, leading to the following mechanical equilibrium equations. \u23a7 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a8 \u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23aa\u23a9 \u222b \u03b11 0 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 + \u222b \u03b12 0 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 = P 2 , if a \u2208 \u03a9I \u222b \u03b11 0 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 + \u222b \u03b71 0 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 + \u222b \u03b12 \u03b71+\u03b72 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 = P 2 , if a \u2208 \u03a9II \u222b \u03b71 0 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 + \u222b \u03b12 \u03b71+\u03b72 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 + \u222b \u03b73 0 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 + \u222b \u03b11 \u03b73+\u03b74 k\u2217Pcos2\u03b8 2r1 r1d\u03b8 = P 2 , if a \u2208 \u03a9III (2) B", " (2) gives the following expression of k\u2217 in each region. \u23a7 \u23aa \u23aa \u23aa \u23aa\u23a8 \u23aa \u23aa \u23aa \u23aa\u23a9 k\u2217 = 4(2\u03b11 + 2\u03b12 + sin2\u03b11 + sin2\u03b12) \u2212 1 , if a \u2208 \u03a9I k\u2217 = 4 [ 2\u03b11 + 2\u03b12 \u2212 2\u03b72 + sin2\u03b11 + sin2\u03b12 +sin2\u03b71 \u2212 sin2(\u03b71 + \u03b72) ]\u2212 1 , if a \u2208 \u03a9II k\u2217 = 4 \u23a1 \u23a3 2\u03b11 + 2\u03b12 \u2212 2\u03b72 \u2212 2\u03b74 + sin2\u03b11 +sin2\u03b12 + sin2\u03b71 + sin2\u03b73 \u2212 sin2(\u03b71 + \u03b72) \u2212 sin2(\u03b73 + \u03b74) \u23a4 \u23a6 \u2212 1 , if a \u2208 \u03a9III (3) In the following, the calculation of the inscribed angle of each arc segment is conducted. Assuming the radius and the central angle of the length arc 1\u20132 are R1 and \u03c61, respectively, \u03b22 shown in Fig. 3(b) is derived as arcsin(l1 /2R1), and then \u03b21 is expressed with \u03b22 as \u03b21 = \u03c61/2 \u2212 \u03b22. Similarly, \u03b23 and \u03b24 can be expressed with the radius and the central angle of the width arc 1\u20134, which are denoted as R2 and \u03c62, respectively. Hence, the angles of \u03b11 and \u03b12 are derived as \u03b11 = \u03be1 + arcsin l1 2R1 , \u03b12 = \u03be2 + arcsin l2 2R2 (4) where \u03be1 = arctan R2sin(\u03c62/2) R1sin(\u03c61/2) \u2212 \u03c61 2 , \u03be2 = arctan R1sin(\u03c61/2) R2sin(\u03c62/2) \u2212 \u03c62 2 (5) In Eq. (5), R1sin(\u03c61/2) is half the distance between corner 1 and corner 2, and R2sin(\u03c62/2) is half the distance between corner 1 and corner 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure20.19-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure20.19-1.png", "caption": "Figure 20.19 Plate clamp for vertical lift", "texts": [ " Collar eyebolts may be used up to the SWL for axial lifting only. Eyebolts with a link (Fig. 20.14) offer considerable advantages over collar eyebolts when loading needs to be applied at angles to the axis (Fig. 20.15). Their SWL is relatively greater than those of the plain collar eyebolt and the load can be applied at any angle. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 20 Moving loads 20 302 Where large plates are to be lifted, plate clamps (Fig. 20.19) are available which are used to vertically lift plate up to 130 mm thick and having a SWL of 30 tonnes. Chains with welded links in alloy steel are manufactured in accordance with British Standards (BS) covering the designation of size, material used and its heat treatment and dimensions, e.g. material diameter, welds and dimensions of links. The welds should show no fissures, notches or similar faults and the finished condition should be clean and free from any coating other than rust preventative" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000377_tsmc.2019.2931740-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000377_tsmc.2019.2931740-Figure1-1.png", "caption": "Fig. 1. LMDIP system.", "texts": [ " Then, we find the approximate solution by three-layer feedforward NN. Finally, a discrete-time NN controller is synthesized for the LMDIP system. In Section IV, we verify the control algorithm by experiment and make some comparisons with polynomial approximation method. In the end, we make some conclusions for this article in Section V. In this section, we will establish the discrete-time model of the LMDIP system and then formulate the position tracking control problem of the LMDIP system as an approximate discrete-time NOR problem. Fig. 1 shows the LMDIP system, where d \u2208 R denotes the position of the cart, Mc denotes the mass of the cart, \u03b2 \u2208 (\u2212\u03c0, \u03c0) denotes the angle between the rod and the vertical line, u \u2208 R denotes the control force, lr denotes the length from the center of mass to the base of the pendulum, Mr denotes the mass of uniform rod instead of block on the pendulum as in [1] and [30], and Ir denotes the moment of inertia of the pendulum. Based on the Lagrangian dynamics, the dynamic model can be deduced as follows [40]: u = (\u03b2\u0308 cos\u03b2 \u2212 \u03b2\u03072 sin\u03b2)Mrlr + d\u0308(Mc + Mr) 0 = \u03b2\u0308(Ir + Mrl2r )\u2212 Mrglr sin\u03b2 + Mrlrd\u0308 cos\u03b2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003757_09544054211028515-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003757_09544054211028515-Figure6-1.png", "caption": "Figure 6. Diagram of the cutting area.", "texts": [ " When the cutting speeds are 30 and 50m/min, TS2500 shows better tool life, while at the cutting speed of 100m/min, TS2000 exhibits better tool life. Figure 5 illustrates the wear distribution on the flank faces of 890, TS2000, and TS2500 inserts when they reach the failure criteria at vc=50m/min. Obviously, the wear concentrates on the region near the tool nose. The average wear width near the minor cutting edge is larger than that near the depth of cut. Diagram of the cutting area is shown in Figure 6, which reveals that under 95 cutting edge angle, 5 end cutting edge angle and 0.5mm depth of cut, the tool nose arc acts as the main cutting edge and the chip thickness varies at different positions of the cutting edge. Continuous friction between major flank face and machining surface induces high temperature and large contact force, being the main reasons for tool failure. When the tool feeds, uncut material exists as shown in Figure 6(c) and it is plowed to the cutting edge in the form of side flow. The friction and extrusion effect of the uncut material and the serious work-hardening effect make the flank wear near the minor cutting edge larger. Figures 7 to 9 present FSEM micrographs of the flank wear at the depth of cut side of the three types of inserts at vc=30, 50, and 100m/min, respectively. Homogeneous wear is observed and the wear mode includes adhesion and abrasive wear. The flank wear of TS2000 at vc=30m/min is shown in Figure 7(b) and (e) and it can be seen that the flank wear could be divided into the following three areas in the perpendicular direction: (1) Area A is the main adhesion area and the area shows more scratches", " According to the EDS analysis of the point 4 (see Figure 10(d)), the elements belong to GH2132, indicating that the pits are produced due to the rupture of BUE since BUE is not stable during the cutting process. Similar phenomenon is also found for 890 at the cutting speed of 100m/min as shown in Figure 15(d). The rupture of the BUE or BUL also takes away some tool material, which leads to the rake face wear. Near the minor cutting edges, adhesion is much serious, in particular, at higher cutting speed as shown in Figures 14 and 15. On the cutting edges near depth of cut, slighter adhesion is observed, while cracks and breakage are distinct. It can be seen from Figure 6 that the shorter the distance to uncut material, the thinner the cutting layer. The thin cutting layer makes adhesion mainly concentrate on the cutting edge. At the depth of cut, the tools suffer serious thermal and mechanical stress. Considering the high hardness and brittleness of carbide tools, the large stress gradient at the boundary leads to cracks and breakage. When the cutting speed is 100m/min, coating peeling is observed for the coated tools as shown in Figure 15(b) and (f). TS2500 even produces thermal cracks in the intact coating area near boundary of BUL" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002328_j.mechmachtheory.2020.103873-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002328_j.mechmachtheory.2020.103873-Figure1-1.png", "caption": "Fig. 1. Process of hypoid gear generating methods.", "texts": [ " This paper focused on the strength comparison between face-milled and face-hobbed hypoid gear. The accurate mathematical model of hypoid gear were established. The finite element analysis models were developed to investigate the root stress and bending stress for both two types of hypoid gears. The standards of strength check were used to evaluate those two types of hypoid gear and verify the results from finite element (FE) analysis. Two different cutting method are involved in hypoid gear manufacturing process. Those two process methods are shown in Fig. 1 ( a ) - ( b ). For face-hobbed process, the tooth trace of gear is extended epicycloid line. The cutter head, gear blank and two cradle (generator cradle and epicycloid cradle) rotation contribute to continuous indexing. The rotation relationships are shown as following ( \u03c9 c2 \u2212 \u03c9 c1 ) / \u03c9 t = z 0 / z p (1) ( \u03c9 c2 \u2212 \u03c9 c1 ) / \u03c9 p = z/ z p (2) For face-milled process, the tooth trace is circular arcs, so the epicyclical cradle rotated angle should be zero ( \u03c9 c1 = 0 ). In order to focus on single indexing (cutting one tooth slot at one moment), the rotation relationship can be obtained by \u03c9 c2 / \u03c9 p = z/ z p (3) where \u03c9 p and \u03c9 t represent the rotated angular velocity of gear blank and cutter head, respectively. \u03c9 c 1 denotes the velocity of epicycloid cradle (first cradle). \u03c9 c 2 is the rotated angular velocity of generated cradle (second cradle). z is the teeth number. z 0 denotes the number of cutter dead and z p represents the teeth number of generated gears. The structure of face-hobbed and face-milled cutter are shown as Fig. 1 . For face-hobbed gear, the cutter head is divided into two groups (inner and outer blade cutter head). Inner blade and outer blade cutter head also contribute to convex and concave gear flank of hypoid gear. Therefore, R ac and R av represent the radius of outer blade and inner blade, respectively. For face-milled gear, the cutter head is divided into two surfaces (inner and outer surface). The R represent the nominal radius of cutter head. The mathematical model of three-face cutter for face-hobbed hypoid gear included transfer matrixes and coordinate system is proposed in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002404_00423114.2020.1752923-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002404_00423114.2020.1752923-Figure1-1.png", "caption": "Figure 1. Simplified model of the hybrid powertrain.", "texts": [ " Based on the use of Taylor expansion at \u03c90 = 0 and the neglect of higher order term, Equation (1) can be rewritten as \u03bc( \u03c9) = \u03bcs \u2212 \u03b4(\u03bcs \u2212 \u03bcd)e\u2212\u03b4| \u03c9|( \u03c9 \u2212 \u03c90)+ \u00b7 \u00b7 \u00b7 = \u03bcs \u2212 \u03b4(\u03bcs \u2212 \u03bcd) \u03c9sgn( \u03c9) (2) Then, the clutch torque can be expressed as Tc = N\u03bcsFcARc ( 1 \u2212 \u03b4 \u03bc \u03bcs \u03c9sgn( \u03c9) ) = Tc0 ( 1 \u2212 \u03b4 \u03bc \u03bcs \u03c91 ) (3) where \u03c91 = \u03c9\u00b7sgn( \u03c9), N is the number of friction pair in clutch, Fc is the compression force of platen,A is the contact area of friction pairs, Rc is equivalent radius of friction pairs, Tc0 is the constant composition of clutch torque and \u00b5 is the difference between static friction coefficient and dynamic friction coefficient, i.e. \u00b5 = \u00b5s \u2212 \u00b5d. In this paper, we take the mode transition process in a PSHEV as the research object, and the simplified diagram of the corresponding hybrid powertrain is shown in Figure 1. When the hybrid powertrain switches from electric mode to hybrid driving mode, the clutch is gradually engaged and transfers the engine torque to the motor to jointly drive the load. The dynamic equation of the hybrid powertrain during mode transition can be written as J\u03c9\u0307 = Tc + Te \u2212 TL \u2212 \u03b2\u03c9 (4) where J is the combined inertia of the clutch driven plate and the motor, Te is the motor output torque, TL is the load torque, \u03b2 is the viscous damping coefficient of clutch driven plate and motor and \u03c9 is the motor angular speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000959_s00502-014-0272-3-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000959_s00502-014-0272-3-Figure3-1.png", "caption": "Fig. 3. Structure of a double-sided axial flux machine with internal rotor [3]", "texts": [ "de); Reis, Kersten, Institut f\u00fcr Elektrische Energiewandlung, Technische Universit\u00e4t Darmstadt, Landgraf-Georg-Stra\u00dfe 4, 64283 Darmstadt, Deutschland; Binder, Andreas, Institut f\u00fcr Elektrische Energiewandlung, Technische Universit\u00e4t Darmstadt, Landgraf-Georg-Stra\u00dfe 4, 64283 Darmstadt, Deutschland 25 types can be designed as slotted or unslotted machines. The unslotted machines have a big magnetic air-gap, as the windings of those machines are placed in the air-gap. So low flux densities occur in the air-gap, and only low forces can be generated with unslotted machines, but there is nearly no torque ripple. This report focuses on the double-sided machines with a slotted stator and a stator tooth coil winding leading to a high power density and a high pole count. The axial force is balanced by both machine sides and is extinguished. Figure 3 shows the double-sided AFM with internal rotor (AFIR). It has two stator cores, which carry the stator winding and one internal rotor disc. The rotor disc consists of permanent magnets, which are arranged as a ring shape. The magnets are carried by a nonmagnetic and non-conductive carrier construction to avoid eddycurrents due to the stator magnetic field. Due to the two stators, no rotor iron yoke is required to guide the flux. The flux passes directly through the PM, which reduces the axial machine length between the two air-gaps" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001588_s12206-015-1118-6-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001588_s12206-015-1118-6-Figure7-1.png", "caption": "Fig. 7. Scheme on the impact hammering modal test.", "texts": [ " The results showed that a loud noise occurred when the nut passed along the 100-110 mm and 130-150 mm sections of the lead screw, as shown in Figs. 6(a) and (b), respectively. In the experiment, friction noise in the lead screw system occurred only at the certain location of the nut and it was reproducible. To verify the relationship between the fundamental frequency of friction noise and the nut location, a hammering test was conducted on the lead screw system. The external forces (F1, F2) were generated by an impact hammer with the variation of the nut location, and the acceleration (aY, aZ) was measured by accelerometer as shown in Fig. 7. To determine the torsion mode, impact torque were generated by impacting the hitting point 'p1' with the tangential force ' F1', and the tangential acceleration 'aZ' were measured. The axial mode was also confirmed by impacting the point 'p2' with the axial force (F2) and measuring the acceleration 'aY' at the opposite end of the lead screw. Fig. 8(a) showed that the natural frequency of 3100 Hz corresponded to the torsion mode of the lead screw and it rarely changed with the nut location. It confirmed that the 3100 Hz friction noise aroused from the torsion mode of the lead screw" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000898_chicc.2015.7260508-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000898_chicc.2015.7260508-Figure2-1.png", "caption": "Fig. 2: Definition of vehicle body-axis, thrust, torque etc.", "texts": [ " In Section 3, a Linear Quadratic Regulator (LQR) based control framework for vehicle attitude angles execution is firstly introduced, based on which position tracking of the Qball-X4 is discussed. Experimental tests for *Professor Youmin Zhang is currently on sabbatical leave from the Department of Mechanical and Industrial Engineering, Concordia University, Montreal, Quebec H3G 1M8, Canada. The object vehicle discussed in this report is built up based on Qball-X4, see Fig. 1. Certain definitions of the vehicle body-axis and angular rotations are presented in Fig. 2. Given that the thrust and torque generated by the 4 electric motors are 1 2 3 4 1 2 3 4, , , , , , ,T T T T respectively, total thrust and rotational torque along each axis are: 1 2 3 4 3 4 1 2 1 2 3 4 ( ) ( ) zu T T T T u L T T u L T T u (1) Note that individual torque generated by the motor is given by ( 1,2,3,4 )i i iK T , and L is the distance between motors and vehicle center. Equation (1) could be written as: 1 2 3 4 1 1 1 1 0 0 0 0 zu T u L L T u L L T u K K K K T (2) Thrust delivered from each motor is controlled by the input voltage" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000868_gt2015-43971-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000868_gt2015-43971-Figure5-1.png", "caption": "Figure 5: A CAD model of an impeller to be manufactured using laser consolidation process.", "texts": [ " The surface finish of the as-consolidated IN-718 flat specimens was measured using a Mitutoya Surftest 402 Surface Roughness Tester. The average roughness (Ra) of asconsolidated IN-718 alloy is about 1.6 \u03bcm along horizontal direction and 2.94 \u03bcm along vertical direction. Building Impellers Laser consolidation is a material addition process that can directly build functional features on an existing component to form integrated structure without the need of welding or brazing. For example, LC can be used to build net-shape functional components on pre-machined substrate. An impeller shape (Figure 5) was selected to demonstrate the capability of LC process using a 5-axis CNC motion system. The impeller has a diameter of about 77 mm and height of about 26 mm. There are 9 long blades and 9 short blades uniformly distributed. A profiled substrate was pre-machined using a CNC lathe (Figure 6) and was mounted on a 5-axis CNC motion system using a simply designed fixture. Using a 5-axis CNC motion system to deal with large tiltrotation movement, LC was successfully conducted to build blades on the pre-machined substrate to form an integrated impeller" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001483_0954406215612814-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001483_0954406215612814-Figure3-1.png", "caption": "Figure 3. Graphical application of the Euler\u2013Savary equation to find both centers of curvature and s of the corresponding conjugate profiles and s.", "texts": [ " In particular, the center of curvature s of profile s coincides with the contact point P, which means that the involute profile s has a cusp at P. In fact, the line attached to the auxiliary centrode \" (line) passes in this case through the center of curvature Ol of the centrode l. This feature will be clarified in the Euler\u2013Savary equation for envelopes and Aronhold theorems and the return circle sections based on the Euler\u2013Savary equation for envelopes and Aronhold\u2019s first theorem, which takes into account the properties of the return circle. The Euler\u2013Savary equation for envelopes can be obtained by referring to the sketches of Figure 3, which show the graphical determination of the centers of curvature and s, of and s, respectively, as envelopes of during the pure-rolling motions of \" on l (Figure 3a) and \" on l (Figure 3b), respectively. Thus, both conjugate profiles and s can be considered in contact at the point P of Figure 4, where the graphical construction of Figure 3 is applied again by showing that the centers of curvature and s are also, correspondingly, the centers of curvature for their relative motion represented by the centrodes l and l, whose centers of curvature are Ol and Ol , respectively. Thus, the Euler\u2013Savary equation for envelopes can be expressed as 1 P0 1 P0 l cos \u00bc 1 rl 1 rl \u00f01\u00de where is a polar coordinate of point P with respect to the ordinate axis of the canonical frame (P0, t, n), P0 and P0 l giving the positions of the centers of curvature and l with respect to P0, while rl and rl are the signed radii of curvature of l and l, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003601_tmag.2021.3081799-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003601_tmag.2021.3081799-Figure1-1.png", "caption": "Fig. 1. Sectional view and trimetric view of 12/12/11 FM-TFLG. (a) Sectional view. (b) Trimetric view.", "texts": [ " Different combinations of permanent magnet number, slot number and translator number are selected and their performances are compared and analyzed. This paper is organized as follows. In Section II, the machine configuration and operation principle are introduced. In Section III, some dimension parameters are optimized to maximize the average thrust. In Section IV, the performance of 12/12/11 and 18/12/17 FM-TFLGs are compared, including open-circuit flux linkage, thrust, back EMF, and efficiency, followed by conclusions in Section V. II. MACHINE CONFIGURATION AND OPERATION PRINCIPLE In the proposed structure, as shown in Fig. 1, the armature winding is placed in outer stator slots. Excitation methods include electric excitation and PM excitation. Direct current (DC) field winding and PM are both located in inner stator. The translator consists of spiral silicon steel sheets and nonmagnetic sheets. In order to facilitate manufacturing, the translator is divided into four segments. Two nonmagnetic ribs are used to assemble the four segments of the translator. The spiral translator of the proposed machine is the only D Authorized licensed use limited to: Carleton University", " The main difference is that, FRPM is a kind of rotating machine and the proposed machine is a transverse flux linear machine. In DD-WEC, due to the incoming wave, the buoy is driven to move up and down, thus the translator of the generator moves along the z-axis. The gear box and other intermediate transmission device are not needed when using linear generator [3]. This means that the whole system is simplified and the efficiency is improved, so the proposed machine is more suitable for DD-WEC. The permanent magnet number, slot number, translator number of the generator in Fig. 1 are 12, 12, 11, respectively. Generally speaking, the 12-stator-pole machine with 11- translator-pole translator has better performance than 12/10, 12/13 and 12/14 counterparts [11]. Such as, higher power density, smaller thrust ripple and better sinusoidity of back EMF. In order to study whether higher modulation ratio can obtain better performance, such as higher power density, better sinusoidity of back EMF and higher efficiency, 12/12/11 and 18/12/17 FM-TFLGs are analyzed and compared in this paper, the translator numbers of both topologies are one less than the PM numbers of these two machines" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000613_j.mechmachtheory.2019.103718-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000613_j.mechmachtheory.2019.103718-Figure10-1.png", "caption": "Fig. 10. Schematic diagram of the 3PRRR PM.", "texts": [ " Similarly, the average resulting angular EDC for each component can be obtained as shown in Table 5 . It is found that the resulting angular deflection of the mechanism is mainly caused by the bottom shaft. It is also noteworthy that the angular stiffness performance of the PTPM cannot be improved by increasing the stiffness performance of the actuators. Finally, the proposed method is used to study the linear and angular stiffness performances of a 3PRRR PM with three translational DOFs, as illustrated schematically in Fig. 10 . The elastostatic analysis of this mechanism was previously addressed in Ref. [24] . However, the EDC of each elastic component of the mechanism was not discussed. The parameters used in the elastostatic analysis are shown in Table 6 . Each limb exerts two constraint couples on the moving platform, oriented perpendicular to the axis of the R-joint of the limb, and one driving force along the direction of the R-joint axis [24] . The compliance matrix for each link corresponding to the constraint wrenches determined via the strain energy method can be expressed as follows: C i 1 = \u23a1 \u23a2 \u23a3 l i 1 G i 1 A i 1 + l 3 i 1 3 E i 1 I i 1 l 2 i 1 2 E i 1 I i 1 0 l 2 i 1 2 E i 1 I i 1 l i 1 E i 1 I i 1 0 0 0 l i 1 G i 1 J i 1 \u23a4 \u23a5 \u23a6 , (51) C i 2 = \u23a1 \u23a2 \u23a2 \u23a3 l i 2 G i A i + l 3 i 2 +3 l 2 i 1 l i 2 c 2 \u03b8i +3 l i 1 l 2 i 2 c \u03b8i 3 E i 2 I i 2 + 3 l 2 i 1 l i 2 s 2 \u03b8i G i 2 J i 2 2 l i 1 l i 2 c 2 \u03b8i + l 2 i 2 c \u03b8i 2 E i 2 I i 2 + l i 1 l i 2 s 2 \u03b8i G i 2 J i 2 2 l i 1 l i 2 s \u03b8i c \u03b8i + l 2 i 2 s \u03b8i 2 E i 2 I i 2 \u2212 l i 1 l i 2 s \u03b8i c \u03b8i G i 2 J i 2 2 l i 1 l i 2 c 2 \u03b8i + l 2 i 2 c \u03b8i 2 E i 2 I i 2 + l i 1 l i 2 s 2 \u03b8i G i 2 J i 2 l i 2 c 2 \u03b8i E i 2 I i 2 + l i 2 s 2 \u03b8i G i 2 J i 2 l i 2 s \u03b8i c \u03b8i E i 2 I i 2 \u2212 l i 2 s \u03b8i c \u03b8i G i 2 J i 2 2 l i 1 l i 2 c \u03b8i s \u03b8i + l 2 i 2 s \u03b8i 2 E i 2 I i 2 \u2212 l i 1 l i 2 c \u03b8i s \u03b8i G i 2 J i 2 l i 2 c \u03b8i s \u03b8i E i 2 I i 2 \u2212 l i 2 c \u03b8i s \u03b8i G i 2 J i 2 l i 2 s 2 \u03b8i E i 2 I i 2 + l i 2 c 2 \u03b8i G i 2 J i 2 \u23a4 \u23a5 \u23a5 \u23a6 , (52) where C i 1 and C i 2 denote the compliance matrices for the links A i B i and B i C i that correspond to the constraint wrenches, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure3-1.png", "caption": "Fig. 3. Dynamic mesh and contact lines of helical gear with geometric eccentricity.", "texts": [ " 2 shows that btan\u03b2b is less than LCD when \u03b5\u03b1 is greater than \u03b5\u03b2 , and btan\u03b2b is greater than LCD when \u03b5\u03b1 is less than \u03b5\u03b2. In the figure, T represents the time required for a single tooth to rotate a tooth distance, and n is the minimum integer greater than the total contact ratio. The transverse contact ratio \u03b5\u03b1 and axial contact ratio \u03b5\u03b2 can be expressed as tan,= =CD b t t L b P P\u03b1 \u03b2 \u03b2\u03b5 \u03b5 , (14) where pt is the normal transverse circular pitch. Moving coordinate system O\u00b4X\u00b4Y\u00b4 is built as shown in Fig. 3. The geometric center O1\u00b4 of driving gear is fixed on the original point O\u00b4 of moving coordinate system. The geometric center of the gear O2\u00b4 is on the X axis. O2\u00b4 moves between point M and point N given that the geometric center distance varies with time t. O1 and O2 are the centers of rotation. The actual length of surface of action CD in Fig. 3 can be obtained from the surface of action truncated by the gear pair addendum cylinder. The whole surface of action is shown as ABEF. L1, L2, L3...Ln represent the contact line of each tooth pair, and n is the smallest integer greater than the total weight of the total contact ratio. In the reference coordinate system O\u00b4X\u00b4Y\u00b4, the position of addendum and base circle centers of the driving gear remain fixed, while the addendum and base circle center positions of the driven gear move between M and N. Point C moves between the dotted lines of C1 and C2 when the position of driven gear and base circle changes, as shown in Fig. 3. Point C0 is the initial position. At t = 0, the tooth only enters meshing, and the length of the contact line at the point C0 is 0. The position of the dotted lines C1 and C2 is the limited position of intersection line between the addendum cylinder of the driven gear and the surface of action. When the geometric center distance becomes larger, the geometric center of driven gear moves toward point N, and point C moves to point C2. As a result, the total length of the contact line is reduced. When geometric center distance becomes smaller, the geometric center of driven gear moves toward point M, and point C moves to the point C1. Accordingly, the total length of the contact line is increased. In Fig. 3, LAC0 represents the distance from the point A to C0 at initial position, and LAC represents the distance from the point A to C at any position. As shown in Fig. 4, a meshing period is divided into four parts according to the different positions of the contact line. In the case of Fig. 4(a), the axial contact ratio is larger than the transverse contact ratio. However, in Fig. 4(b), the transverse contact ratio is larger than the axial contact ratio. The length of the meshing line varies periodically" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure8.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0001681_ccdc.2014.6852993-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001681_ccdc.2014.6852993-Figure6-1.png", "caption": "Figure 6: Shear force", "texts": [ " Lxi+1\u2212xi\u22121 represents the distance between the position of xi+1 and the position of node xi\u22121. 2l represents the relax distance between the position of xi+1 and the position of node xi\u22121. Uxi+1\u2212xi\u22121 represents the vector obtained by subtracting the position of xi+1 from the position of xi\u22121. The force of Fa.pring connects two specified nodes with an angle spring, as shown in Fig. 5. Fshear(i) =\u2212kshear(\u03b8i \u2212\u03b8b) \u00b7u (7) where kshear represents shear force coefficient. The shear force straightens the catheter/guide wire tip and keeps the original shape of the tip, as shown in Fig. 6. The internal forces applied on the body are Fspring, Fshear, and Fbending. For ith particle, it is defined as follow: Fspring(i) = Fs.spring(i) (8) Fbending(i) = \u2212kbendingcross((xi \u2212 xi\u22121) ,( Ux \u2016Ux \u2016 )) \u00b7 kadd \u00b7 \u2016 Fadd \u2016 (9) where kbending represents bending coefficient. The function of cross is defined as the cross-product operation (the crossproduct is given by the vector obtained by subtracting the position of xi from origin and the vector obtained by subtracting the position of xi\u22121 from origin) " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003598_tmag.2021.3078841-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003598_tmag.2021.3078841-Figure1-1.png", "caption": "Fig. 1. Configuration of the proposed IMR-AMGDRM. (a) Overall view of configuration. (b) Exploded view of configuration.", "texts": [ " The accuracy and the cause of calculation error of the analytical model are investigated. II. CONFIGURATION OF THE IMR-AMGDRM In the conventional AMGDRM, the inner and outer rotors are the PM and the modulating rotors, respectively. Since the modulating rotor is closer to the PM rotor than stator, it suffers from high axial magnetic force. Furthermore, the modulating rotor is hollow and thus low mechanical strength becomes one of the major drawbacks [1]. To solve this problem, the IMR-AMGDRM is proposed, of which the configuration is shown is Fig. 1. The IMRAMGDRM consists of three parts: stator, the PM rotor and the modulating rotor. The modulating rotor is composed of ferromagnetic pole-pieces and back iron, which means that it can be manufactured as one solid core and thus the mechanical structure is strengthened. Moreover, the IMR-AMGDRM can be viewed as the one transforming from the conventional AMGDRM by switching positions of the PM and the modulating rotors. Therefore, in the IMR-AMGDRM, the inner and outer rotors are the modulating and the PM rotors, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001323_ihmsc.2014.92-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001323_ihmsc.2014.92-Figure5-1.png", "caption": "Figure 5. Left graph is radial pattern arrangement, right graph is regular polygon formed by 8 robots.", "texts": [ " With the growth of the aggregation, if the robot joined aggregation stops immediately, all the robots arrange in multi-tree (Fig.4). In order to increase perception range of aggregation, this paper proposes that the robots arrange in radial pattern or regular polygon. The radial pattern arrangement refers to when robot joins aggregation, it will adjust its own position to make the robots arrange in line. When the line reaches the edge of arena, robots will arrange in new line whose starting point is the halfway point of the original straight line (Fig.5). The arrangement of regular polygon refers to the robots in aggregation arrange in regular polygon by local positioning and information interaction. The distance among contiguous robots is no more than the sensing range of each robot in any arrangement. Assuming a uniform distribution of robots in the bounded arena, constant speed, and constant sensing range, any robot in searching state has a constant probability pc to encounter another individual of constant size and appearance at every time step of length T" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003746_s10409-021-01089-9-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003746_s10409-021-01089-9-Figure8-1.png", "caption": "Fig. 8 von Mises Stress distribution of the beam. a Solid-beam element result. b Ansys result", "texts": [ " In the first mesh, 7 elements are used in the transverse direction, while in the second mesh there are 28. At the longitudinal direction, 6, 12, 24 and 36 elements are used respectively. This example is also performed by the ANCF 96-DOF solid element with the same element mesh as validation. Comparative results are exhibited in Fig.\u00a07. As a benchmark, the same problem is solved in ANSYS. 588,336 solid45 elements are used in total and results indicate that the free tip displacement is 2.3089 m. In Fig.\u00a08, the distributions of von Mises stress are compared between the proposed solid-beam element and Ansys. From the results one can find that the 96-DOF solid element exhibit higher precision than the proposed solidbeam element. However, the solid element mesh has as twice degree-of-freedom and nine times number of integration points as the solid beam element mesh. It is known that the evaluation of the elastic force and its Jacobian is one of the most time-consuming part in the ANCF computer implementation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001998_msec2016-8516-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001998_msec2016-8516-Figure4-1.png", "caption": "Figure 4: (a): 3D Point cloud of NIST part (100 mm \u00d7 100 mm \u00d7 8 mm; 4 inch \u00d7 4 inch \u00d7 0.3 inch) obtained using the FaroArm laser scanning probe. (b): Zoomed in section showing fine features (4x zoom).", "texts": [ "org/about-asme/terms-of-use practically tenable, two different sampling methods are pro- posed to the SGT approach in Sec. 3.2.3. (3) Application of SGT to AM parts and verification with artificially generated surfaces (Sec. 4): The above twostep SGT approach is applied to AM test parts (Sec. 4.1), and subsequently verified with artificially generated test parts (Sec. 4.2). The test artifact printed for this study was designed and de- veloped by Moylan et al. [29] at NIST. The so-called NIST test artifact has become the industry standard for comparison among various AM processes and material combinations. Figure 4 illustrates the various part features, including cylinders, holes, ramps, fine features, edges, staircases, pins, holes, and flat surfaces. The part measures ~ 100 mm \u00d7 100 mm \u00d7 8 mm (4 inch \u00d7 4 inch \u00d7 0.3 inch). Each feature is designed to evaluate a specific capability of the AM process based on the limits of dimensional accuracy [29]. Manufacture of Test Components The components examined in this study were manufactured using a polymer-based material extrusion AM process, called fused filament fabrication (FFF) [3]", " For instance, the first component (Figure 3(a) is made with acrylonitrile butadiene styrene (ABS) plastic in a machine having a thermally controlled build chamber, and is thus labeled as ABS chamber. Table 1 summarizes the conditions for the manufacture of the three different types of AM components, namely, ABS chamber, ABS platform, and CF-ABS platform. Description of Measurement Procedure and Data The three samples were scanned with a FaroArm Platinum linear scanning laser probe to generate a three-dimensional point cloud. A representative sample scan consisting of ~ 500,000 data points is shown in Figure 4. The laser scanner records reflected light from the surface of a component as a point in 3D space, with a maximum volumetric deviation of \u00b1 43 \u03bcm. The point cloud for each part was then imported into a commercial software package (Geomagic by 3DSystems) for analysis. Standard functions within the software are used to remove outlier points and disconnected components. The error-corrected 3D point cloud data is subsequently converted to a polygon mesh for comparison against the reference CAD model in order to assess the geometric accuracy of the component", " However, in this method, the windowing procedure is slightly modified. Instead of fixing the window size k as done in SGT-Method 1, in SGT-Method 2 the number of windows \ud835\udcc3 is fixed and then we compute the window size k. This has the effect of slicing the component into \ud835\udcc3 strips of identical width. An example of such spatial sampling is shown in Figure 10(a). If \ud835\udcc3 is set at 500, as is done for all instances in this work, then each sampling window is approximately 10 mils in width (250 \u03bcm), and 0.3 inch (8 mm) in height for the NIST sample (Figure 4). Because of such spatial sampling, the sequence \ud835\udeb22(\ud835\udcc3) = [\u03bb2 1 \u03bb2 2 \u22ef \u03bb2 \ud835\udcc3]T is mapped to a particular area on the component. Consequently, it is possible to track which facet has deviated along with the magnitude of the deviations from the design blueprint dimensions. The SGT approach is now applied, first to the experimentally acquired point cloud data (Sec. 4.1), followed by analysis of artificially generated samples (Sec. 4.2). These numerical studies augment the experimental results by seeking to quantify the change in Fiedler number with controlled variation in dimensional integrity" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000629_iecon.2019.8927706-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000629_iecon.2019.8927706-Figure1-1.png", "caption": "Fig. 1. Different housing types for TEFC method [3].", "texts": [ " However, higher current density leads to higher copper losses and higher temperature. Therefore, to achieve the required power and performance, the thermal analysis of the electrical machine is essential. Different applications of electrical machines require different cooling methods. The electrical machine faces several thermal constraints. These thermal drawbacks are modified and moderated by selecting an appropriate cooling method. Accordingly, the industrial motors in the medium power range are cooled by the totally enclosed fan cooled (TEFC) method. As figure 1 shows, in this cooling approach, the outer circumferential of the machine housing consists of the fin channels and the fan blades are mounted in the nondriven section of the machine shaft to provide the air flow into the fin channels during rotation of the motor shaft. The objective of this paper is to enhance knowledge in thermal design and analysis of the TEFC electrical machine and consider different thermal analysis tools for electrical machines with a particular focus on SynRMs. The research focuses on existing challenges and problems in thermal analysis of TEFC machines" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003485_j.jmatprotec.2021.117165-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003485_j.jmatprotec.2021.117165-Figure3-1.png", "caption": "Fig. 3. (A) Schematic diagram of S-DED system. (B) Assembly of system integrated with the cooling channel and scanning DED nozzle.", "texts": [ " With the help of this numerical simulation, it was confirmed that the angle of inlet of the nozzle (where the metal powder flows first) significantly affects the difference in the amount of powder finally captured by the laser beam, and the difference in powder height of each position. Based on the results of the simulation, an angle of inlet of 75\u25e6, a 5- channel outlet, and a 3-channel cooling section were selected for the optimized design of the S-DED nozzle. In real case, we choose 5-channel outlet to prevent clogging of metal powder on the outlet entrance. Fig. 3 depicts the model of the S-DED system designed on the basis of the numerical simulation results. The angle of the inlet through which a mixture of gas (N2) and powder flows is 75\u25e6. The nozzle consists of a 5- channel outlet and three cooling channels through which the metal powder is cooled. It has a sealing pocket for shielding the gas and powder, and a powder storage through which the metal powder passes through inside the nozzle. Fig. 3(A) shows the S-DED head combined with the high-throughput nozzle. The main head is further combined with two nozzles and a protection cap. The nozzle has a nozzle plate for sealing; a shielding gas (Ar) is continuously sprayed from the protection cap to prevent damage to the laser lens, due to fumes and spatter. The S-DED system, developed according to the modeling results, is shown in Fig. 3(B). The protection cap is made of aluminum; and the scanning nozzle, nozzle plate and main head are made of copper for fast cooling because copper has good thermal conductivity. Thus, to summarize, the S-DED system consisted of a TUEMPF 2 kW diode laser, a FANUC 6-axis robot, and a platform with an 8 \u00d7 3 mm square flat-top beam shown in Fig. 4. In the experiment, white metal of ASTM B 23 grade 2 was used, and the size of the powder was 25\u2013125 \u03bcm. When different values of hopper speed was used (1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001428_1350650114559617-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001428_1350650114559617-Figure3-1.png", "caption": "Figure 3. Roller contact with inner ring and outer ring.", "texts": [ " Before the analysis, the following assumptions are made: (1) the deformation is in elastic range; (2) the shaft fitted with inner ring has no elastic deformation; (3) the bearing house fitted with outer ring has no elastic deformation; (4) geometrical error was only about roller diameter; (5) if the roller has some geometrical error, with the load application, the roller gets rigid body motion; (6) bearing is 0 internal radial clearance. Based on the above assumptions, deformation occurs only in the area in which roller contacts with the inner ring and the outer ring, as shown in Figure 3. In Figure 3, O and O1 are the center of inner ring and the roller respectively before loading; O0 and O01 are respectively the center of inner ring and the roller after loading. When the roller is squeezed, it will produce a deformation between the roller and inner ring and outer ring. The elastic deformations of the roller between inner and outer ring are i and o respectively and has the relationship R R0 \u00bc i \u00fe o \u00f09\u00de Assuming the roller load is Q, we can obtain i \u00bc 0:2723E0L cos i Dmr 0:074 \" # 1 1:074 Q 1 1:074 \u00f010\u00de o \u00bc 0:2784E0L 1\u00fe o t 0:078 \" # 1 1:078 Q 1 1:078 \u00f011\u00de Therefore, the total elastic deformation of roller with inner ring and outer ring can be expressed as \u00bc i \u00fe o \u00f012\u00de \u00bc 0:2723E0L cos i Dmr 0:074 \" # 1 1:074 Q 1 1:074 \u00fe 0:2784E0L 1\u00fe o t 0:078 \" # 1 1:078 Q 1 1:078 \u00f013\u00de Figure 4 shows that the rolling elements are in the asymmetric state" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002775_j.triboint.2020.106669-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002775_j.triboint.2020.106669-Figure6-1.png", "caption": "Fig. 6. The finite element model of the tribo-system, (a) components of the model, (b) load and boundary conditions of the model.", "texts": [ " This analysis procedure is commonly used for solving nonlinear problems without considering the iteration and convergence criteria. It is based on the implementation of an explicit centraldifference time integration rule together with the use of diagonal lumped element mass matrices. It can evaluate the normal and tangential contact stresses along the contact region [30,31]. Thus, the friction effect between the roller and ring can then be fully considered in the analysis. A simplified finite element model, as shown in Fig. 6, is created according to the geometry of the experimental system. Fig. 6(a) shows the components of the finite element model. The ring is fixed to the supporting shaft and the roller shaft is attached to the drive shaft. The two supporting bearings are simplified as two ferrules with their upper parts cut into planes and tied with the load block. Based on the experimental measurement, a friction coefficient between the roller and ring is set to 0.2. The roller shaft and the supporting bearings are assumed to be in friction contact with a friction coefficient of 0.01", " Material properties of the parts used in the numerical calculation are based on their actual values, as listed in Table 3 Material properties of the parts in finite element model. Part Density (kg/ m3) Young\u2019s modulus (GPa) Poisson\u2019s ratio Drive shaft 7820 210 0.288 Roller 7800 207 0.3 Ring 7800 207 0.3 Load block 7850 210 0.31 Supporting bearing 7800 207 0.3 Supporting shaft 7850 210 0.31 C. Xu et al. Tribology International 154 (2021) 106669 Table 3. The load and boundary conditions of the model are shown in Fig. 6 (b). The rotational velocity boundary conditions are applied to the supporting shaft and the drive shaft in the x-direction. The normal load is applied to the block in the y-direction. Other constraint conditions are all consistent with the real experimental system. In the dynamic analysis, the element type used is C3D8R, the contact formulation is the kinematic method with the finite sliding formulation and the friction formulation is the penalty method. The total simulation duration is 0.502 s" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000623_0954406219893721-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000623_0954406219893721-Figure4-1.png", "caption": "Figure 4. Tooth profile of involute gear and novel HCR gear: (a) pinion tooth and (b) internal gear tooth.", "texts": [ " With the development of computer hardware and FE software, it is now feasible to work out these rather complex models. The process of creating involute gear pair and novel HCR gear pair is completed with the design language of MATLAB and geometric modeling software CATIA. The first step is the generation of enough number of tooth profile key points coordinates by the tooth profile mathematical model equations, following which is the build of high precision 2D planar model. The comparison of tooth profile between the involute internal gear and the novel HCR internal gear is shown in Figure 4 according to the geometry parameters in Table 1. Take the consideration of the spur gear and the short computation time, the 2D FE model is well suitable for the simulation of a variety of different gear pairs. A precise geometric model is a good basis for FE simulation. What\u2019s more, a relatively fine mesh density is also critical to simulate the non-linear contact deformation faster and better between the tooth surfaces. Kiekbusch et al.12 introduced two methods to refine the mesh in the area of contact during the simulation process" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002260_j.engstruct.2020.110257-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002260_j.engstruct.2020.110257-Figure9-1.png", "caption": "Fig. 9. Failure mode and force-displacement curve for node in tension.", "texts": [ " The maximum forces applied to each bolt are 5kN and 2.63kN respectively for tension node and compression node, and the maximum bending moment applied to each branch of the bending node is 0.13kNm. The failure modes of tension, compression and bending nodes obtained from numerical models are shown in Figs. 9\u201311. The forcedisplacement curves measured at a reference point of each node are also shown in Figs. 9\u201311. As can be seen, the failure of the tension node occurs in the short members of the node as circled in Fig. 9. The failure in bending node only takes place in planar substructures in compression side of the node due to buckling, while the other side under combined bending and tension is not failed. Based on the predicted maximum loads that the structural node can sustain in different loading conditions, the loads required in the experimental tests can be calculated, considering the deterministic structural system of the test rig. Details of the applied loads and the dimensions of test rig in different tests are shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure9.32-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure9.32-1.png", "caption": "Figure 9.32 Top slide set at half included angle", "texts": [ " If a deeper form is required, the knurls are released and moved back to the start before picking up in the original indents, increasing the force on the knurls and the operation repeated. Figure 9.29 Diamond knurl Figure 9.30 Basic knurling tool This method is normally used for short tapers such as chamfers, both internal and external. The long cutting edge required by long tapers has a tendency to chatter, producing a bad surface finish. Taper turning can be carried out from the top slide by swivelling it to half the included angle required on the work, Fig.\u00a09.32. Graduations are provided on the base plate, but any accurate angle must be determined by trial and error. To do this, set the top slide by means of the graduations, take a trial cut, and measure the angle. Adjust if necessary, D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 9 141 9 Turning The guide block slides on the guide bar and is located in the sliding block by a spigot. This gives a solid location and at the same time allows the guide block to take up the angle of the guide bar" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000717_s00542-015-2728-8-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000717_s00542-015-2728-8-Figure1-1.png", "caption": "Fig. 1 Illustration of top and end plate structures", "texts": [ " This study aims to fabricate a self-pumping glucose oxidase fuel cell, delivering fuel into interval cell through capillary force between electrode and flow field without an extra pump. Compared to existing active systems, the cell scale is much smaller, and working without an extra pump which is more cost-effective. The experimental process in this study includes the selfpumping fuel cell structure design and fabrication. The MFC glucose oxidase (GOx) solution and immobilization is prepared. Then cell is assembled and tested before experimental analysis. The top plate structure is shown in Fig. 1. There are three reservoirs, Reservoir 1 is the fuel supply. In order to make a self-pumping structure, Reservoir 2 and 3 were installed wet cotton. The plates were carved using a CNC machine. There are 27 channels on the bipolar plates. The middle channel is the widest channel which is 800 \u03bcm. The channel widths on both sides decrease arithmetic by 15 \u03bcm, as shown in Fig. 2. The Gas flow channel is set up for gas emissions to deal with CO2 produced on anode side. The field pattern was firstly designed and fabricated in a 1 3 photoresist pattern" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001914_1.b35750-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001914_1.b35750-Figure5-1.png", "caption": "Fig. 5 Mesh of straight-through labyrinth seal.", "texts": [], "surrounding_texts": [ "The general dynamic equation is defined as\n\u2212 fx fy\nKxx Kxy Kyx Kyy\nX Y\nCxx Cxy Cyx Cyy\n_X _Y\n(1)\nwhere [C] and [K] represent the damping and stiffness matrices, respectively; and ffg is an external force vector. In this study, the rotor material density is 7800 kg\u2215m3, the modulus of elasticity is 2.1 \u00d7 1011, the Poisson\u2019s ratio is 0.3, the bearing stiffness is 1.75 \u00d7 107 N\u2215m in every direction, and the damper of the bearing is considered to be zero. After a dynamic analysis, the amplitude and frequency of vibration are input into a CFD computation as unsteady-state parameters. The motion trajectory can be represented as\nx A cos \u03c9t y A sin \u03c9t\n(2)\nwhere the amplitude of vibration is represented by A, and the vibration frequency and time are expressed as \u03c9 and t, respectively. The ANSYS code can calculate the rotor movement accurately using Eq. (2). Flow in the labyrinth seal is three-dimensional; therefore, the three-dimensional Reynolds-averaged Navier\u2013Stokes equations have to be solved by CFD computation to analyze the flow patterns. A commercial code, termed ANSYS CFX\u2122, which uses the finite volume method, can be used to deal with the complex CFD computations. To simulate turbulent flow in the clearance of the labyrinth seal, a suitable turbulence model must be chosen carefully. The realizable k-\u03c9 model is selected in the CFD calculation, and a second-order discrete form with higher-order resolution is picked. In the numerical simulation, the average static pressures on the flowfield outlet are always 0.1MPa; and the total airflow inlet pressures are 0.5, 0.4, and 0.3 MPa. The details are shown in Table 1. In the code, rotor vibration is achieved using moving mesh technology. The motion commands move the nodes on the rotating wall and transmit the movement trend to the adjoining nodes. Under the influence of mesh motion, deformation on the rotating boundary is transmitted as follows:\n\u2207 \u00b7 \u0393disp\u2207\u03b4 0 (3)\nwhere \u03b4 is the displacementwith respect to the originalmesh position, and \u0393disp represents the stiffness of the grid. When the unsteady-state flow computation (including the rotor vibration) has converged, the formula will be solved at the beginning of each time step. In the code, mesh movement can be commanded by setting its stiffness. A growth pattern of the mesh stiffness, which is selected before simulation, \u201cincreases in the small volume.\u201d The stiffness coefficient is varied by\n\u0393disp 1\n\u2200\nCstiff\n(4)\nwhere \u2200 is the control volume size; and Cstiff is the index of stiffness model,whichcanbe set to command the stiffness coefficient of themesh.\nTo analyze the effect of vibration on the labyrinth seal characteristics, the authors define eight uniform points on the movement trajectory of the rotor (Fig. 4). Points 1 to 8 are abbreviated as P1 to P8, and the number sequence is along the circumferential direction of the circle motion. The eight monitoring points constitute a complete vibration cycle, and the calculation data can be analyzed to investigate the effect of seal flow and aerodynamic force because of vibration. Clearance on the object viewed ismagnified to describe the relative position of the rotor movement. In this research, the aerodynamic effect inside the seals is of special concern. Therefore, the aerodynamic force acting on the rotor will be analyzed for discussing the aerodynamic effect and its influence on rotor. Compared with the perturbation method-based CFD method, the method used in this research decomposes the seal aerodynamic force into fx and fy, as well as the radial force and tangential force, instead of using them to solve the four characteristic coefficients. Thus, the relationship between the flowfield variation and rotor position can be described intuitively.\nTo investigate the unsteady-state phenomena in labyrinth seals and avoid the influence on the periodic boundary because of mesh movement, the authors used a complete circle of mesh to replace the periodic mesh in the CFD simulation. The multiblock structured grids for computation were generated using ICEM CFD\u2122. In the meshing process, special locations of the computational model were refined. Figures 5 and 6 show the computational meshes, which were refined near the wall and small boundary. The boundary-layer thickness on the wall reached 5 \u00d7 10\u22126 m, which met the y requirements of the turbulence model. The structure of the interlaced seal was more complex than the straight-through seal, and it required a higher mesh quality. To ensure the precision and dependability of the numerical process for analyzing the flow characteristics of the two types of labyrinth seals, a high-accuracy mesh must be built. Time-consuming computations should also be taken into account. Therefore, the sensitivity verification for the undetermined meshes was processed, and the results are depicted in Fig. 7. The variations of mesh quantity in each direction are proportional, so the total mesh quantity reflects the mesh quality in each direction. After comparison, we obtain a mesh with 2,469,000 elements for the final computational mesh of the straight-through seal and a mesh with 3,516,000 elements for the interlaced seal.\nFigure 8 shows the flow domain and boundary conditions of the CFD simulation. In Fig. 8b, the axial section is cut out from the flow\nD ow\nnl oa\nde d\nby W\nE ST\nE R\nN M\nIC H\nIG A\nN U\nN IV\nE R\nSI T\nY o\nn A\nug us\nt 5 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /1\n.B 35\n75 0", "domain. The boundary conditions of the inlet, outlet, and rotating wall are marked in the chart, and the uncharted boundaries are all no*slip walls. On account of the complete cycle model used in the computation, the periodicity boundary is not adopted in this research.\nAfter the rotor dynamics computation, the natural frequency of the rotor is solved as 247.3 Hz; that is, the critical speed of rotation is 14,838 rpm. After that, the correspondence between the vibration frequency and the amplitude can be acquired. The chart in Fig. 9 depicts the vibration trajectory, which is magnified for easy observation and computed by the rotor dynamics code. From the results, the amplitudes of any node on the rotor axis at different\nrotational speeds can be obtained. In this research, the largest amplitude is chosen as the computational condition for the unsteadystate flowfield simulation. In the process of rotor vibration, if the frequency of vibration does not arrive at the first-order inherent frequency, the amplitude increases with the augment of the frequency of vibration. And, when the frequency of vibration is reached between the first- and secondorder inherent frequencies, the amplitude will increase with the augmentation of the vibration frequency. The designated rotating speeds in this project are greater than the first-order critical speed of rotation (but have not yet reached the second-order critical speed of rotation); therefore, the amplitude of vibration is reduced with the enhancement of the vibration frequency. The interrelation of the\nFig. 6 Mesh of interlaced labyrinth seal.\n0 1500 3000 4500 6000\n0.010\n0.012\n0.014\n0.016\n0.018\n0.020\n0.008\nGrids (thousands)\nM as\ns fl\now (\nkg /s\n)\nStraight-through seal Adopted for computation\nAdopted for computationInterlaced seal\nFig. 7 Sensitivity verification the computational meshes.\nD ow\nnl oa\nde d\nby W\nE ST\nE R\nN M\nIC H\nIG A\nN U\nN IV\nE R\nSI T\nY o\nn A\nug us\nt 5 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /1\n.B 35\n75 0", "rotating speeds, frequencies of vibration, and amplitudes of vibration are shown in Table 2.\nThe lengths of the computational domain inlet and outlet have a great influence on the results of the flow simulation. To improving the accuracy of the CFD computation, the sizes of the inlet and outlet\ncomputational domains must be set appropriately. Particular to the annular seal, which is a type of clearance seal, the flow condition of the passageway near the seal tip is much different from that in the seal cavities. The complete computational domain of the flowfield is constituted by some quick constricted and outspread portions. Thus, on condition that the lengths of the computational domain are insufficient, the results will be unreliable because of the unsteady flow. The nonuniform progress of the outlet flow influences the accuracy of several significant flow parameters, such as the mass flow. Because of the aforementioned considerations, an appropriate computational domain size must be confirmed before formal calculations. And, the effect on the inlet and outlet flowfield duct lengths is illustrated in Fig. 10. Three computational models of the same labyrinth seal are used in the contrast calculation. Threemodels are solved at the same conditions. From the charts, it can be seen that vortices exist in the seal cavities. Vortices in the outlet channels are also visible because of the throttling action of the seal teeth.When the outlet length is 2.9 mm, the outlet channel is filled completely by a vortex, and no smooth axial streamlines exist near the outlet. When the outlet length is 8.9 mm, the outlet channel is filled mostly with a vortex, and the flowfield near the outlet has a smooth axial trend. When the outlet length is 15.9 mm, streamlines near the outlet are mostly smooth, and the outlet flowfield can be considered to be uniform. The actual length of the outlet used in this research is greater than 15.9 mm, so the effect on the outlet length is reduced considerably. To verify the effect by the outlet section length in a quantitative way, we did a sensitivity test in terms of leakage. The verification computation was done under a steady condition by using a periodicity geometric model. The pressure difference in the verification computation was 0.3 MPa. And, the results are shown in Table 3. From the table, it can be seen that the leakages were nearly constant when the outlet length was larger than 15.9 mm. So, the outlet length selected in this research was applicable. A comparison of the leakage at different conditions (Table 4) shows that the rotor vibration increases the labyrinth seal leakage slightly. And, the larger amplitude results in more leakage. That is, rotor vibration plays a role in reducing seal performance. Although the cross-sectional area of the leakage path is not changed, the clearance is not uniform when the rotor is offcenter. Leakage flow will be biased toward the larger clearance in the annular seals, which may increase the seal leakage.\nInlet length 3.3 mm, outlet length 2.9 mm\nInlet OutletSeal section\nInlet OutletSeal section\nInlet OutletSeal section\nInlet length 5.3 mm, outlet length 8.9 mm\nInlet length 10.3 mm, outlet length 16.9 mm\nFig. 10 Influence of inlet and outlet duct length on flowfield.\nTable 3 Relationship between leakage and outlet length\nOutlet length, mm Mass flow, kg\u2215s 2.9 0.00116878 8.9 0.00116286 11.9 0.00116284 14.9 0.00116270 15.9 0.00116266 17.9 0.00116265 20.9 0.00116265\nTable 4 Comparison of mass flow\nRotational speed \u03c9, rpm Amplitude A, mm Pressure difference \u0394p, MPa\nMass flow _m (kg\u2215s) Straightthrough\nlabyrinth seal\nInterlaced labyrinth\nseal\n18,000 0.0927 0.2 0.01745 0.01205 0.3 0.02135 0.01481 0.4 0.02464 0.01710 24,000 0.07945 0.2 0.01727 0.01199 0.3 0.02115 0.01474 0.4 0.02441 0.01703 30,000 0.06275 0.2 0.01711 0.01196 0.3 0.02100 0.01468 0.4 0.02425 0.01697 30,000 0 0.2 0.01670 0.01180 0.3 0.02045 0.01445 0.4 0.02358 0.01663\nD ow\nnl oa\nde d\nby W\nE ST\nE R\nN M\nIC H\nIG A\nN U\nN IV\nE R\nSI T\nY o\nn A\nug us\nt 5 , 2\n01 6\n| h ttp\n:// ar\nc. ai\naa .o\nrg |\nD O\nI: 1\n0. 25\n14 /1\n.B 35\n75 0" ] }, { "image_filename": "designv11_22_0003683_iemdc47953.2021.9449537-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003683_iemdc47953.2021.9449537-Figure11-1.png", "caption": "Fig. 11 Equal potential distributions and flux contours. (a) SPM. (b) CPM1. (c) CPM2. (d) CPM3. (e) CPM4.", "texts": [ " In addition, under overload condition, the output torque of CPM4 is higher than that of CPM1, which mainly results from the larger equivalent air-gap length due to the pole shapes. Under the rated current, the losses of three analyzed machines are compared in TABLE IV. Clearly, CPM1, CPM2 and CPM3 exhibit similar losses, while CPM4 shows the lowest value for both core loss and PM eddy current loss. The decrease of the dominant core loss is because of the smaller high order harmonics and the lower saturation levels in stator and rotor cores as shown in Fig. 11. Thicker permanent magnets of CPM1, CPM3 and CPM4 increase the capability to resist the armature flux and thus lead to lower PM eddy current loss. V. CONCLUSION To achieve the largest output torque and reduce the large torque ripple in CP machines, the rotor PM and iron pole shapes were optimized by GA in this paper. It has been demonstrated that different PM and iron pole shapes and approximately equal PM pole arc span and pole pitch are necessary to achieve both the largest output torque and low torque ripple for a fixed amount of magnet" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000197_00450618.2019.1609088-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000197_00450618.2019.1609088-Figure3-1.png", "caption": "Figure 3. FDM process.18", "texts": [ "5 Three-dimensional printers can be categorized by the different additive printing methods used, as summarized in Table 1. Fused disposition modelling (FDM), also referred to as Fused Filament Fabrication (FFF), is the most widely used processes in personal 3D printers.16,17 Thermoplastic filaments consisting of polymers such as acrylonitrile butadiene styrene (ABS) and polylactic acid (PLA) are passed through a heating block and extruded from a nozzle onto a printer bed. The process is repeated until the finished object is obtained.17 Figure 3 illustrates the FDM process. Stereolithography apparatus (SLA) is the basis of original 3D printers, in which a thin layer of curable liquid photopolymer resin is mechanically spread over the print bed, and a cross-section of the 3D object is cured with a UV laser. The print bed descends, and the process is repeated.19 Digital Light Processing (DLP) printers use UV photopolymerized resins that are fused through a flash process, but unlike SLA printers, where each layer is fused by the laser, this process fuses the entire layer in one process using a highresolution UV light projector" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002534_012146-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002534_012146-Figure7-1.png", "caption": "Figure 7. Distribution of the magnetic field of cross-section of the Stack 2", "texts": [ " To visualize the flow paths of the main magnetic flux: consisting of the sum of the fluxes of permanent magnets and the axial magnetic flux of the superconducting field coil, the simulation results are shown in figures 5 through 13. The figures 5-13 shows that the value of magnetic induction in ferromagnetic parts (tooth, yoke) does not exceed the permissible values. The figures 5 through 8 show cross sections of an electric machine in the plane of the first packet (Fig. 5), in the plane of the axial interpacket yoke (Fig. 6), the second packet (Fig. 7), and also between the second packet and the axial excitation coil (Fig. 8). 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10.1088/1742-6596/1559/1/012146 An analysis of the magnitude of the magnetic induction shows that the iron is not in saturation mode - the magnitude of the induction does not exceed 2.1 T. The direction of magnetization of the poles of 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10", "1088/1742-6596/1559/1/012146 permanent magnets and ferromagnetic poles is alternating. The values of the fluxes of pole scattering and scattering by the antennae of the stator teeth are insignificant. An analysis of the magnitude and direction of the axial magnetic flux in the sectional plane of the interpacket yoke (Fig. 6) shows that the fluxes from the poles in it are alternating. This confirms the need to use twisted silicon electrical steel as an interpacket yoke material to reduce the magnetic losses in it. Figure 7 shows that the threephase coils of the superconducting windings of the armature of the second package are displaced by 15 geometric degrees relative to the coils of the first package. The constancy of the magnitude and direction of the axial component of the induction in the yoke of the bearing shield and in the ferromagnetic shaft show that there is no need to use a special electrotechnical steel as magnetic circuit materials. 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001472_s0263574714002458-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001472_s0263574714002458-Figure5-1.png", "caption": "Fig. 5. Some examples for cases considered in the cross-section collision check.", "texts": [ " Algorithm 3 SS-CollisionCheck(seci, f ) /* This algorithm checks for collisions between a cylinder arm section and a polygon */ Compute the distance between straight-line segment segi and the plane Q that the object face f is on, denote the distance as d(segi, Q) and closest points p, q on segi and Q, respectively if d(segi, Q) \u2264 wi then if q is not on f then compute the minimum distance between each edge ek of f and segi ,41 denoted as dmin(segi, ek) with closest points p and q on segi and ek respectively if dmin(segi, ek) > wi then return Collision \u2190 False end if end if if p is not on pi\u22121 or pi then return Collision \u2190 True else /* f is closest to an end point of segi */ if f intersects Hi or Hi\u22121 at lint and the distance d(lint , pi) \u2264 wi then return Collision \u2190 True end if end if end if return Collision \u2190 False. fan-shaped cross-section csi of manipulator section i. Denote the polar coordinates of v1 and v2 as (\u03c11, \u03b81) and (\u03c12, \u03b82), respectively. Algorithm 4 classifies all collision scenarios into five cases based on whether v1 and v2 satisfy the bounds of inequalities (1) and (2) and then check sequentially from Case 1 to Case 5 to detect all possible kinds of intersections (i.e., collisions). Figure 5 shows examples for these cases. Appendix A further shows that Algorithm 4 covers all cases of possible collisions and is complete. If a face f of the object does not intersect Pi or does not intersect the cross-section csi , we need to further check if f intersects section i by Algorithm 5. In Algorithm 5, we first check if the distance between ciri and Q, the supporting plane of face f , is greater than the width of section i by calling Procedure 1. If so, Q has no intersection with the section i, a truncated torus, and no further collision checking is necessary", " Since N n (as it requires more than one bounding volume for each arm section), and n1 n, MN + mn1 (M + m)n. Thus, the CD-CoM Algorithm has a far lower order of worst-case time complexity than OPCODE, and we can further show that O(MN + mn1) O[(M + m)n]. https:/www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574714002458 Downloaded from https:/www.cambridge.org/core. Georgetown University Library, on 13 Feb 2017 at 18:46:01, subject to the Cambridge Core terms of use, available at Arm model Algorithm Config. 4 Config. 5 Config. 6 Exact CD-CoM <1 ms 6 ms 10 ms Mesh 1 OPCODE 7 ms 11 ms 16 ms Mesh 2 OPCODE 37 ms 47 ms 53 ms For checking collision of a path of k arm configurations, the worst-case time complexity of CD-CoM is O[k(M + m)n]; whereas, the worst-case time complexity of OPCODE is O[k(MN + mn1) + Nk], with the additional term Nk reflecting the time complexity of refitting the bounding volume hierarchy of the manipulator for each change of configuration. Thus, k(MN + mn1) + Nk k(M + m)n, and moreover, O[k(MN + mn1) + Nk] O[k(M + m)n]", ", min(\u03c11, \u03c12) > ri + wi and \u03b81, \u03b82 both satisfy (or both do not satisfy) (2) then compute the distance between circle center ci and li and obtain point q = (\u03c1q, \u03b8q) on li if the distance is shorter than ri + wi and \u03b8q satisfies (2) return Collision \u2190True else return Collision \u2190 False Case 4: if both \u03b81 and \u03b82 satisfy (2) then return Collision \u2190 True / * the remaining collision cases have intersections between rays Li,1 (or Li,2) and li * / Case 5: if line segment li intersects rays Li,k (k = 1, 2) of csi at pk int (see Fig. 4) then if one vertex of li satisfies (2) and is above the upper bound for \u03c1. then if \u03c1k int \u2264 ri + wi then return Collision \u2190 True (see Fig. 5(e)) end if end if if one vertex of li satisfies (2) and is below the lower bound for \u03c1. then if \u03c1k int \u2265 ri \u2212 wi then return Collision \u2190 True (see Fig. 5(f)) end if end if if neither vertices of li satisfies (2) then if one of \u03c1k int satisfies \u03c1k int \u2265 ri \u2212 wi then return Collision \u2190 True (see Fig. 5(g)) end if end if return Collision \u2190 False. https:/www.cambridge.org/core/terms. https://doi.org/10.1017/S0263574714002458 Downloaded from https:/www.cambridge.org/core. Georgetown University Library, on 13 Feb 2017 at 18:46:01, subject to the Cambridge Core terms of use, available at Algorithm 5 NCS-Collision check(seci , f ) /* This algorithm checks for collisions between a toroidal arm section seci and a polygon f when f does not intersect the section plane Pi */ Compute the minimum distance between ciri and the plane Q of face f from the object, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002403_s11277-020-07378-z-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002403_s11277-020-07378-z-Figure2-1.png", "caption": "Fig. 2 Body-fixed coordinate system [20]", "texts": [ " The planned system falls in the category of actuated system due to the equivalent number of DOF and thrusters. Figure\u00a01, shows the overall model of the underwater vehicle that is equipped with a camcorder. Table\u00a0 1, defines the position (x, y, z) of ROUV and roll, pitch and yaw ( , , ) respectively which describes the orientation of body. Linear velocity components are (u, v,w) and angular velocity components are (p, q, r) of ROUV. The forces and movements of the body are (X, Y , Z) and (K,M,N) respectively. Figure\u00a02, shows the orientation of body-fixed coordinate system that are used for standard motion of ROUV [17]. Figure\u00a02, defines the earths fixed frame and the body fixed coordinate system. The proposed ROUV consist of six thrusters and individually of one gives the notation f1\u22366 for the force provided by them. Table\u00a03 defines the movement and Table\u00a04 defines the turning on and off of thrusters respectively. The physical parameters of the vehicle are defined in Tables\u00a02 and\u00a03 demonstrate the notations for the thrusters of a 6-DOF of ROUV. Table\u00a0 4 the directions produced by these thrusters. ( T1 \u2212 T6 ) are the thrusters located on ROUV", " From the past, different researchers design ROUV, by using different structures which are based on diverse number of thrusters to made it stable. Consequently, many of them unsuccessful because of the high pressure and some are failed because of not proper maneuvering in deep sea. In this study, our design ROUV armed with six thrusters, in which four of them are placed at the bend of the ROUV. Fifth and sixth are positioned in the center of the body of our underwater vehicle in the upward and downward direction which is shown in Fig.\u00a02. Model reference adaptive control (MRAC) along with the integral action use for the feedback loop regularly use in unmanned systems. Moreover, PID controller is responsible for the fine tuning of the gains of the system. On the other hand, the dynamic system stability is dealt by the flexible control laws of an adaptive controller. Remark 1 Now, the unknown system non-linearity\u2019s and model uncertainties are not included in the system model [ As(t),Bs(t) ] . The uncertainty co-efficient \u201c \u201d is present in the specification of hydrodynamic" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002975_sensors47125.2020.9278807-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002975_sensors47125.2020.9278807-Figure1-1.png", "caption": "Fig. 1: A schematic of the bevel-tipped needle in the ideal insertion into two-layered tissue, which is modeled as two different uniformly distributed loads (normal components are only shown).", "texts": [], "surrounding_texts": [ "Index Terms\u2014flexible needle; FBG; mathematical model; tissue inhomogeneity; needle motion planning;\nI. INTRODUCTION\nSince the pioneering work of unicycle/bicycle model [1], [2], there have been many good models applied to flexible needle steering. The different kinds of models include kinematics-based models [1]\u2013[3] and mechanics-based models [4]\u2013[6]. Kinematics-based models have been used in control and planning of the needle, but may not properly capture additional or complex needle deformation. Mechanicsbased models are usually based on classical beam theory to determine needle shape, but require detailed information on the tissue. During prostate surgery, for example, accurate needle placement through insertion into multi-layer tissue is required to prohibit from significant tissue damage. A sensorbased model using a Lie-group theoretic approach [7] has been proven successful for shape-sensing in multi-layer tissue using FBG sensors [8]. This model is based on the theory of elastic rods and Lie groups with the advantage of complex needle shape recognition in 3D space through the use of FBG sensors.\nAside from needle shape-sensing, an important issue to transition into human tissue is the prediction of needle shape during insertion in multi-layer tissue. Prediction of needle insertion is directly performed in kinematics-based models by integrating the system equation, however, the model that we apply treats the needle as a loaded elastic rod, requiring special consideration. In this paper, we propose and evaluate our needle insertion prediction model, expanding upon [7], [8], for single- and double-layer insertion scenarios. Specifically, we investigate necessary factors of the model that are influenced by increasing insertion lengths. To do that, we model uniformly-distributed loading conditions over increased\nThis work has been supported by the National Institutes of Health under grant No. R01CA235134.\nlengths of an elastic rod in order to determine proper scaling parameters in needle curvatures. Amidst the COVID-19 pandemic, we rely upon previous data presented in [7] to perform our prediction optimization in real data for single- and doublelayer insertion scenarios. In Section II and III, the model and the simulations are explained. After the simulation results in Section IV, we conclude the paper.\nThe sensor-based model (see [7] for the details) describes the curvature (\u03c91 and \u03c92 along the local x- and y-axes, respectively) and torsion (\u03c93 along the local z-axis) of the needle as\n\u03c9(s) = [\u03c91 \u03c92 \u03c93] T = ( RT (s)dR(s)\nds\n)\u2228 (1)\nwhere R(s) \u2208 SO(3) denotes the rotation matrix in 3D space describing the orientation of the body-fixed frame attached at each point along the needle. s \u2208 [0, L] denotes the arclength of the needle with total insertion length L. Here \u2228 operation defines a 3D vector associated with a 3 \u00d7 3 skew-symmetric matrix RT dRds [9].\nWhen inserting into tissue, the needle is modeled as an inextensible elastic rod. Under ideal conditions, we assume that the needle experiences uniformly distributed loads. In case of homogeneous tissue (i.e., single-layer) insertion, a single uniformly-distributed load and for two-layered tissue, two different, uniformly-distributed loads are assumed. In all cases, due to an asymmetric bevel tip, the needle deforms only in one plane (yz plane in the world frame), which is captured by introducing the intrinsic curvature \u03ba0(s). For the single- and double-layer cases, they both introduce \u03bac constants, denoted as intrinsic curvature constants. Refer to [7] for details. Then, through the minimization of the elastic potential energy in the\n978-1-7281-6801-2/20/$31.00 \u00a92020 IEEE\nAuthorized licensed use limited to: Makerere University Library. Downloaded on May 14,2021 at 08:33:24 UTC from IEEE Xplore. Restrictions apply.", "rod using the Euler-Poincare\u0301 (E-P) equation [10], [11] together with FBG sensor data, we determine the deformation, \u03c9(s), and subsequently, r(s), the points along the needle.\nThe optimization variables, intrinsic curvature constants and the initial rod deformation, are constant over the length of the needle in shape-sensing. In determining these variables, FBG sensors are indispensable. To predict the shape of the needle further inserted, we must investigate the dependence of the optimization variables on varying insertion lengths. For practical purposes, given a determined shape at a reference insertion length Lref , we aim to predict the shape at a given insertion length L. In order to provide a simple, yet effective approach for accurate predictions, we propose the following form for predicting \u03bac for increased insertion lengths\n\u03bac(L) = \u03bac,ref (Lref/L) p , (2)\nwhere \u03bac,ref is the optimized \u03bac parameter at the insertion length Lref and p is the scaling parameter to be optimized. For flexibility in multi-layer insertion scenarios, we propose using (2) for all \u03bac parameters associated with each layer using a single, optimized value of p. This claim is verified by modeling distributed loading forces on an inextensible elastic rod for varying lengths as shown in Section III-A. Then with the similar scaling form of the initial deformation, \u03c9init, needle shape prediction is performed, as explained in Section III-B.\nNote that the intrinsic curvature includes the external effects for the bending of the needle. Hence in order to investigate the intrinsic curvature constant, we need to solve the equations for an inextensible rod. This rod theory [12] has been successfully applied to cables and double-stranded DNA [13]. Here we assume the needle insertion as a quasi-static process. Necessary rod equations are written as\nd\nds (B\u03c9) + \u03c9 \u00d7 (B\u03c9) + e3 \u00d7 f = 0 (3)\ndf ds + \u03c9 \u00d7 f = \u2212fext (4)\nwhere B denotes the stiffness matrix. f(s) and fext(s) respectively denote the internal force distribution and external force distribution along the needle. For an ideal needle insertion into homogeneous phantom tissue (single bend), fext = [0,\u2212fn,\u2212ft]T where fn and ft denotes the normal and tangential (or frictional) force densities, respectively. Then we solve (3) and (4) by optimization with the cost function F = \u2016\u03c9(L)\u20162+ \u2016f(L)\u20162+ \u2016\u03c9(0)\u2212B\u22121mt\u20162+ \u2016f(0)\u2212 ft\u20162 where mt and ft respectively denote the total moment and force applied to the rod. Each term corresponds to a boundary condition: zero moments and forces at the distal end, and balance of moments and forces at the proximal end.\nIn solving the rod equations (3) and (4) using the distributed loading tangential and normal forces, ft and fn, respectively we proposed using a scaling of distributed force densities as\nfn = fn,ref (Lref/L) \u03bd ; ft = ft,ref (Lref/L) \u03bd , (5)\nwhere Lref is the reference length to be modeled, and \u03bd is an optimized parameter. The primary assumption associated with ideal insertion of a flexible needle is that the shape from s \u2208 [0, Lref ] for the predicted shape will be approximately the same as the reference shape (i.e. the needle follows the path created from its previous insertion). For two-layer cases, we consider (5) for individual layers, requiring \u03bd1 and \u03bd2 to optimize.\nIn order to generate valid predicted insertions, we propose the following optimization problem. Given a series of rod shapes associated with insertion lengths, we define a cost, Carea, as the mean pair-wise surface area between rod shapes. Secondly, we define a cost on tip deviation Ctip, that is the mean of the pair-wise tip deviation between rod shapes. Finally, the cost function to be minimized is C = Carea+Ctip where we optimize \u03bd in the force scaling. The cost function is intended to minimize the deviation for further insertion of a flexible needle, as per our assumption of ideal needle insertion. Note that this cost function is used for all optimizations in this paper, including the value p in (2).\nFor ideal insertions, we have that the initial condition for integrating the E-P equation, \u03c9init = \u03c90 = [\u03ba0, 0, 0] T , would require a scaling of the same form as (2) with the same value of p. However, in order to account for real-world insertion experiments, we propose a similar, but individual scaling of \u03c9init:\n\u03c9init(L) = \u03c9init,ref (Lref/L) q (6)\nwhere q is a non-universal, optimized parameter. To allow for complex variation in shape, we use q as a separate, optimized parameter per Lref as opposed to p, predicted to be a single, universal value. Finally, with the scaled initial rod deformation, needle shape prediction is performed simply by integrating the E-P equation with scaled \u03bac\u2019s using a single parameter p.\nDue to limited access to experimental research during the COVID-19 pandemic, we utilize previous single- and doublelayer insertion data taken from [7] for needle shape prediction. We simulate needle shapes of increasing insertion lengths and minimize the cost C for the set of shapes through optimizing q in (6). Upon generating the optimized shapes, we perform a workspace analysis through the rotation of the angular deviation vectors, \u03c9 and \u03c90, around its body-fixed z-axis in (1). For q optimization, we simulated shapes from the reference length in our data, 90 mm, up to 150 mm in increments of 15 mm.\nFirst, we perform a modeling of the uniformly-distributed loading on an elastic rod with increasing lengths considered, as\nAuthorized licensed use limited to: Makerere University Library. Downloaded on May 14,2021 at 08:33:24 UTC from IEEE Xplore. Restrictions apply." ] }, { "image_filename": "designv11_22_0002459_978-3-030-16718-9_2-Figure2.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002459_978-3-030-16718-9_2-Figure2.5-1.png", "caption": "Fig. 2.5 Basic (single) supercapacitor cell", "texts": [ " The manufacturing will be formed with evacuation with final welding and sealing. In order to match different applications, ESs should be connected in series or in parallel to obtain the desired output voltage or energy capacity. 20 2 Electrochemical Supercapacitors: History, Types \u2026 This is the last step in which a supercapacitor is manufactured. The rating and applications, the parameters which can vary commercial products and research devices, are strongly affecting factors on the design, assembly, and packaging of the supercapacitor. Figure 2.5 shows details of a basic single supercapacitor cell. Commercially available supercapacitors often prefer stacked cylindrical or coin form. Each manufacturer has its own propriety packaging systems. Cylindrical, typically called \u2018jelly rolls\u2019 because of their shape require the formation of the electrode layer film by rolling or spraying a carbon material on both sides of a separator. An outer separator is applied to ensure the layers are electrically isolated from each other. While preparing coin cell assembly, the cells get pressed and sealed to achieve good electrical contacts and prevent electrolyte leakage, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001211_2582051.2582101-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001211_2582051.2582101-Figure5-1.png", "caption": "Figure 5. System configuration in TAMA II.", "texts": [ " The output component consists of a compact inflator (MARUNI Industry) for bicycles, a CO2 cartridge (BARBIERI) and a DC motor (SCL16-30: NAMIKI PRECISION JEWEL) to open or close the valve of the jet. The input component consists of a 6DoF IMU (SEN-10121: Sparkfun) and a potentiometer to measure the rotation of the motor to control the quantity of ejected gas. In addition, this ball can receive commands from the PC by the Bluetooth module. The total weight of this prototype is 378g, including16g consumable CO2 cartridge. We have succeeded in reducing the weight of 249g from TAMA I. Figure 5 shows the system configuration in the prototype. Table 1 shows the basic features of TAMA I and TAMA II. The Figure 6 shows the preliminary setup to measure the ejection force of the gas, simply putting a pressure sensor seat in front of the nozzle. In TAMA II, it was lighter than previous ones by installation of the new compressed CO2 jet system but the jet pressure decreased. However, the most important requirement is to change the ball\u2019s trajectory. From this point of view, impulse-weight ratio should be the most important factor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001412_j.apm.2015.08.008-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001412_j.apm.2015.08.008-Figure1-1.png", "caption": "Fig. 1. Schematic illustration and geometrical conditions of the EBDMR process with horizontal feeding of the raw material. G1 \u2013 the top surface of the formed pure ingot; G2 \u2013 the interface molten ingot/water-cooled crucible side wall; G3 \u2013 the interface ingot/vacuum; G4 \u2013 the interface ingot/water-cooled puller; wall \u2013 the width of the crucible side wall, under \u2013 the width of the water-cooled puller; Q \u2013 the height of the heat contact at G2.", "texts": [ " In this study, the physical model [53] is treated by an economic and conservative locally one-dimensional numerical method in which the influence of the thermo-physical properties of the metal is more precisely taken into account. Another advantage of this method is that it can easily be continued to a three dimensional method. The numerical model and the optimization problems are developed to analyze and compare experiments and numerical data and to aid in understanding and optimizing e-beam melting. The model of heat transfer, describing the evolution of the temperature field in a cylindrical metal ingot (Fig. 1), is twodimensional due to the cylindrical symmetry of the sample and four interfaces in the ingot (Fig. 1 - G1, G2, G3, and G4) are considered. Recall the domain of the problem \u0304 = {(r, z, t)|(r, z) \u2208 G\u0304, 0 \u2264 t \u2264 F} = G\u0304 \u00d7 [0, F ], (1) where G\u0304 is the domain in which the space variables r, z are defined: G\u0304 = {(r, z)|0 \u2264 r \u2264 R, 0 \u2264 z \u2264 H}. (2) The sets on which the boundary equations [53] are defined are: G1 = {(r, H)|0 \u2264 r \u2264 R}, G2 = {(R, z)|H \u2212 Q \u2264 z \u2264 H}, G3 = {(R, z)|0 \u2264 z < H \u2212 Q}, G4 = {(r, 0)|0 \u2264 r \u2264 R}. (3) With T(r, z, t) we denote the temperature of the investigated metal in the point with polar radius r, height z in the moment of time t" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002460_042016-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002460_042016-Figure2-1.png", "caption": "Figure 2. General view of the experimental prototype of the mini planner.", "texts": [ " Below in figure 1 the location of the screws in the bucket of the planner is showed. However, such improvement in the work of the screw working element for our case, as shown by selective experiments with the experimental sample of a mini-planner, occurs up to the speed of 2 m/s of ICMSIT 2020 Journal of Physics: Conference Series 1515 (2020) 042016 IOP Publishing doi:10.1088/1742-6596/1515/4/042016 translational movement of the unit. Above this speed, the screws begin to be clogged with soil and the technological process of the screw working body is violated. Below in figure 2, experimental sample of a mini-planner with a screw working body is showed. The above analysis requires investigating the productivity of the screw working body, depending on the rotation speed, diameter and pitch of the screw. Therein, the translational speed of the planning unit is also of significant importance. Because, the volume of the planner\u2019s bucket filled with soil per unit of time should be equal to the volume of the processed soil by the screws Below are the curves (figure 3, 4 and 5) of the change in productivity of the screw working element of the planner depending on the rotation speed, diameter and pitch of the screw" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000043_s12541-019-00047-7-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000043_s12541-019-00047-7-Figure3-1.png", "caption": "Fig. 3 Coordinates of the skew rack-cutter: a parabolic profile of the rack-cutter in the normal section; b for derivation of the rack-cutter surface", "texts": [ " First, we use a bi-cubic B-spline to fit the modification curve of the pinion shown in Fig.\u00a02. Next, the fitting surface is superimposed on the theoretical tooth surface to form the modified tooth surface, whose position vector and normal vector are expressed as follows: where Rt (u,\u00a0l) and nt (u,\u00a0l) are the position vector and normal vector of the theoretical tooth surface; \u03b4 (u,\u00a0l) is the modification value obtained from the fitting surface; and u and l are the surface parameters of the skew rack-cutter shown in Fig.\u00a03. The generation of the tooth surface is based on the enveloping of the rack-cutter surface. In Fig.\u00a03, M is a point on the surface of the rack-cutter, whose coordinate is (u, \u2212 acu2, 0) in the coordinate system Sb; dp is the location parameter of the parabolic pole; and \u03b2 is the helix angle. According to the meshing principle, the contact mark of the tooth surface and the gear transmission error can be (1)Rm(u, l) = (u, l) \u22c5 nt(u, l) + Rt(u, l), (2) m(u, l) = m(u, l) u \u00d7 m(u, l) l = ( t(u, l) u + \ufffd(u, l) u \u22c5 t(u, l) + t(u, l) u \u22c5 \ufffd(u, l) ) \u00d7 ( t(u, l) l + \ufffd(u, l) l \u22c5 t(u, l) + t(u, l) l \u22c5 \ufffd(u, l) ) , (a) 2", " 11 Th e te et h 39 40 41 42 43 44 45 46 47 Pi tc h Er ro r ( \u03bcm ) \u2212 0. 89 \u2212 2. 67 0. 89 0 \u2212 2. 22 0. 44 0 0 2. 22 1 3 calculated in the meshing process of a pair of teeth, shown in Fig.\u00a04. Scholars have extensively studied LTCA, while research considering SRNPE is rare. We introduce SRNPE into LTCA by the following steps: First, we calculate the value of SRNPE using Eq.\u00a0(3). where fpb is SRNPE, whose value is shown in Fig.\u00a05; fp1 and fp2 are the measured pitch errors of the pinion and the gear shown in Fig.\u00a03; \u03b2 is the helix angle; and \u03b1n is the normal pressure angle. The tooth numbers of the pinion and the gear are co-prime, so the SRNPE circulates on n (n = 19 \u00d7 47 = 893) meshing periods, and we do not need to consider the initial phase of the pinion and the gear when they enter meshing. Second, we should use TCA stated in Sect.\u00a02.1.1 to get the initial gap between engaging teeth. (3)fpb = ( fp2 \u2212 fp1 ) \u22c5 cos \u22c5 cos n Third, we add SRNPE of the mating teeth to the corresponding initial gap and do 893 loops of LTCA" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002534_012146-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002534_012146-Figure13-1.png", "caption": "Figure 13. Isometric view of the distribution of the magnetic field", "texts": [ " Figures 9 and 10 in the ZX plane. Figure 9 shows the normal component of induction, and figure 10 shows the axial component of induction Bx. In figures 11 and 12 in the section plane YZ, respectively. It can be seen from the figures that the induction is distributed evenly and the axial magnetic circuits are unsaturated, which confirms the correctness of the analytical calculations. The general distribution of magnetic induction and the closure path of magnetic fluxes are shown in an isometric view (Figure 13). The detailed distribution of magnetic induction from Figures 9 to 12 and the general distribution of magnetic field in figure 13 are necessary to understand the principle of operation of an electric machine. Traditional induction machines with an external closed magnetic circuit have significant mass characteristics. The mass of the presented Induction superconducting machine with combined excitation is significantly reduced due to an increase in linear load and magnetic induction in the air gap. And also due to the provision of an alternating magnetic field in the air gap. The mass diagram is shown in figure 14. 14th European Conference on Applied Superconductivity (EUCAS 2019) Journal of Physics: Conference Series 1559 (2020) 012146 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000496_aim.2019.8868817-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000496_aim.2019.8868817-Figure1-1.png", "caption": "Fig. 1. The structure of the SIAT-4 exoskeleton robot.", "texts": [ " Firstly, this paper briefly introduced the active control flow mentioned above and the fourth generation of lower-limb exoskeleton robot developed by Shenzhen Institutes of Advanced Technology of Chinese Academy of Sciences. Then the processing of sEMG signals, the prediction of joint angles and active control of exoskeleton were introduced in detail. Finally, the control method was verified by online experiments. The SIAT-4 exoskeleton robot is the fourth generation lower extremity exoskeleton robot developed by the team of Shenzhen Institutes of Advanced Technology (SIAT), Chinese Academy of Sciences (CAS). The SIAT-4, of which the structure is shown in Figure 1, can help patients with lower limb dysfunction to perform rehabilitation training or assist their walking. With more optimal structure and more wearing comfort, the SIAT-4 mainly consists of mechanical body, control system, drive system and sensing system, weighing only 15Kg. The mechanical body consists of hip, knee, ankle joints, connecting rods, backpack, controller hardware and batteries. Among them, the controller hardware and battery are placed in the backpack; the length of the thigh and calf of the exoskeleton is adjustable, and can be adapted to the wearer with a height ranging from 150cm to 185cm" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001851_s12206-016-0614-7-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001851_s12206-016-0614-7-Figure4-1.png", "caption": "Fig. 4. Timoshenko beam.", "texts": [ " The information about shackle, eyehook, bushings, and clips, which only belong to the traditional leaf spring, are required when MSC.ADAMS/CHASSIS leaf spring is used to develop the taper leaf spring model of tandem suspension [13]; thus, the general parts should be deleted in the taper leaf spring model of tandem suspension. Every piece of leaf spring in the computational model using MSC.ADAMS/CHASSIS leaf spring is divided into several elements because the leaf spring is a continuous flexible body, and each discrete element can be regarded as rigid body and connected by Timoshenko beam, as shown in Fig. 4. The Ti- moshenko beam creates a linear translational and rotational force between two locations that define the endpoints. Eq. (1) shows that the force and torque it applies depends on the displacement and velocity of the I marker on the action body relative to the J marker on the reaction body. The constitutive Eq. (1) is analogous to those in the finite element method as follows: 11 22 26 33 35 44 53 55 62 66 11 12 13 14 15 16 21 22 23 24 25 26 31 32 33 34 35 36 41 42 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 x y z x y z KF x L K KF y K KF z KT a K KT b K KT c C C C C C C C C C C C C C C C C C C C C C -\u00e9 \u00f9\u00e9 \u00f9 \u00e9 \u00f9 \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa = - -\u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00ea \u00fa \u00ea \u00fa\u00eb \u00fb\u00eb \u00fb \u00eb \u00fb 43 44 45 46 51 52 53 54 55 56 61 62 63 64 65 66 x y z x y z V V V C C C C C C C C C C C C C C C w w w \u00e9 \u00f9 \u00e9 \u00f9 \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa \u00ea \u00fa\u00eb \u00fb\u00eb \u00fb (1) where Fx, Fy, and Fz are the measure numbers of the translational force components in the coordinate system of the J marker; x, y, and z are the translational displacements of the I marker with respect to the J marker measured in the coordinate system of the J marker; L is the instantaneous displacement vector from the J marker to the I marker; Vx, Vy, and Vz are the time derivatives of x, y and z, respectively; Tx, Ty, and Tz are the rotational force components in the coordinate system of the J marker; a, b, and c are the relative rotational displacements of the I marker with respect to the J marker as expressed in the x-, y-, and z-axis of the J marker, respectively; and wx, wy, and wz are the measure numbers of the angular velocity of the I marker with respect to the J marker" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001383_smll.201303045-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001383_smll.201303045-Figure2-1.png", "caption": "Figure 2. Transforming a sphere into an ellipsoid by rotated evaporation. a) Schematic geometry showing thickness profi le, h(\u03b8) , of evaporated material that converts sphere to ellipsoid; b) Schematic of rotating evaporation setup; c) Thickness profi les of material deposited onto spheres for various incident evaporation angles \u03b8 i . Theoretically required thickness profi les for creating ellipsoids (dashed curves) closely match those produced by rotated evaporation when \u03b8 i = 30\u00b0 and when \u03b8 i = 70\u00b0.", "texts": [ " With the cavity comprising one ellipsoid subtracted from another, in the limit of fully saturated ferro- magnetic ellipsoids, the generated cavity fi eld becomes equal to the difference between internal fi elds that would have resulted for the outer and inner ellipsoids individually. Uni- form ellipsoid fi elds therefore imply uniform cavity fi elds. In particular, when saturated, ellipsoid magnetizations are equal and cancel to leave an ellipsoid cavity fi eld contribution equal simply to the difference in magnetic depolarization, or demagnetization, of outer and inner ellipsoids. Demagnetiza- tion factors for prolate and oblate ellipsoids magnetized parallel to their axes of revolution ( z -axis in Figure 2 a), are: [ 13 ] 1 1 1 ln 1 1prolate 2 2 2( )= \u2212 \u2212 + \u2212 \u2212 \u23a1 \u23a3 \u23a2 \u23a4 \u23a6 \u23a5D k k k k k (1a) 1 1 1 1 arcsin k 1 oblate 2 2 2 2 = \u2212 \u2212 \u2212 \u2212\u239b \u239d\u239c \u239e \u23a0\u239f \u23a1 \u23a3 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 D k k k k (1b) where k is the ratio of ellipsoid major to minor axes, related to ellipsoid eccentricity through e = [1-k \u22122 ] 1/2 . Note that the demagnetization is independent of overall size. Therefore, for the fi eld within a magnetically saturated ellipsoidal cavity to differ from the applied fi eld, the cavity\u2019s inner and outer ellipsoidal boundaries should have different eccentricities", " KGaA, Weinheim Additionally, to render the cavity accessible, either the shell material should be porous, or the inner and outer bounda- ries should overlap at some point, leading to an opening in the shell. This is possible with either coincident or offset ellipsoid centers. Given their uniform demagnetizations even asymmetric ellipsoidal cavities (as suggested in Figure 1 ) will still generate uniform fi elds. Moreover, this fi eld uniformity can be substantially preserved even for widely open ellip- tical shells. This initially counterintuitive result is possible because overlapped ellipsoids of locally similar curvatures (as suggested in Figure 2 ) can yield large openings with only minimal deviations in shell material thicknesses from the mathematically ideal limiting cases. Although ellipsoidal microparticles are commonly made by stretching [ 14\u201316 ] or irradiating [ 17 ] microspheres, to pro- duce open ellipsoidal microcavities with differing inner and outer eccentricities we introduce a new fabrication scheme. The scheme relies on the observation that, for particular inci- dent angles, evaporation onto a rotating sphere can yield an almost mathematically exact ellipsoid", " Subsequent removal of the spherical core leaves an ellipsoidal microcavity with inner and outer boundaries defi ned by a sphere (zero-eccen- tricity ellipsoid) and a fi nite-eccentricity ellipsoid of revolu- tion, respectively. Evaporation onto spheres is not new. Using such deposition, various two-sided Janus particles [ 18 ] and plasmonically active hemispherical shells [ 19,20 ] have been demonstrated. However, there has been no focus on how shell profi les might be microengineered to generate uniform electromagnetic fi elds within the shell cavity. Transforming a sphere into an eccentric ellipsoid requires particular deposited material thickness profi les. Figure 2 a shows an xz -plane through the middle of a sphere of radius R positioned within an ellipsoid of revolution that has semi-axes R \u00b7 (1+\u03b5 a ) and R \u00b7 (1+\u03b5 b ) and that has been offset vertically by a distance \u03b4 . The crosssectional circle and ellipse are described by x 2 + z 2 = R 2 and [x/(1+\u03b5 a )] 2 + [(z-\u03b4)/(1+\u03b5 b )] 2 = R 2 , respectively, with \u03b4 = \u03b5 b R if circle and ellipse bases coincide as suggested in the fi gure. In polar coordinates this translates into a circle parameterized by r circle (\u03b8) = R and an ellipse by r ellipse (\u03b8) determined from [r ellipse (\u03b8)sin\u03b8/(1+\u03b5 a )] 2 + [(r ellipse (\u03b8)cos\u03b8 -\u03b4)/(1+\u03b5 b )] 2 = R 2 ", " For a thin material shell with \u03b5 a and \u03b5 b much less than unity, this yields the required thickness profi le cos sin cos2 2\u03b8 \u03b4 \u03b8 \u03b5 \u03b8 \u03b5 \u03b8( ) = \u22c5 + +h R Ra b (2) To fi rst approximation, the thickness of material evapo- rated onto a surface scales with the projection of the incident fl uence on that surface. For spherical surfaces this typically yields cosine-like thickness profi les that poorly match the desired h (\u03b8) profi les. However, if material is obliquely evaporated onto a sphere while the supporting substrate rotates about its surface normal (see Figure 2 b), good approximations to h (\u03b8) become possible. Consider an arbitrary point, P , on the surface of such a rotating sphere centered at the origin. By rotational symmetry P can be chosen without loss of generality to have azimuthal angle \u03c6 = 0 , giving position and surface normal vectors proportional to {sin\u03b8, 0, cos\u03b8} . Provided there is direct line of sight to the evaporative source, the instantaneous thickness of material deposited normal to the surface at point P is proportional to the dot product {sin\u03b8, 0, cos\u03b8} \u00b7 {sin\u03b8 i cos\u03c6 i , sin\u03b8 i sin\u03c6 i , cos\u03b8 i }", " KGaA, Weinheim cos cos 2 cos cos cos cos arccos cot cot 2 < < 2 0 2 evap 0 0 \u03b8 \u03b8 \u03b8 \u03b8 \u03c0 \u03b8 \u03c0 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03b8 \u03c0 \u03b8 \u03b8 \u03c0 \u03b8 \u03b8 \u03c0 \u03b8 ( ) ( ) ( ) ( ) ( ) ( ) ( ) = \u2264 \u2212 = \u2212 + \u2212 + \u22c5 \u2212 \u23a1 \u23a3 \u23a2 \u23a2 \u23a4 \u23a6 \u23a5 \u23a5 \u2212 + = \u2265 + h h h i i i i i i i i i (4) The evaporation angle, \u03b8 i , and ellipsoid parameters, \u03b5 a and \u03b5 b , can be connected: for coincident sphere and ellipsoid bases, Equation ( 2) gives \u03b5 a = h (\u03c0/2)/ R and \u03b5 b = h(0)/2 R , yielding, through Equation ( 4) , \u03b5 a = ( h 0 /\u03c0 R )\u00b7sin \u03b8 i , \u03b5 b = ( h 0 /2 R )\u00b7cos \u03b8 i , or, more generally, \u03b5 a /\u03b5 b = (2/\u03c0)\u00b7tan \u03b8 i . Profi les for various \u03b8 i are shown in Figure 2 c. For normally incident evaporation Equation ( 4) reduces, as it must, to h evap ( \u03b8 ) = h 0 cos\u03b8 . For \u03b8 i \u2248 30\u00b0 and \u03b8 i \u2248 70\u00b0 \u2013 75\u00b0, however, the resulting evaporation profi les closely follow the required profi les of Equation ( 2) for generating prolate or oblate ellipsoids with \u03b5 a /\u03b5 b equal to (2/\u03c0)tan 30\u00b0 and (2/\u03c0)tan 70\u00b0, respectively. [ 21 ] The theoretically required and the evaporated profi les overlap so closely that residual mismatches are better gauged by calculating the fi elds that such evaporatively deposited structures would produce when magnetically saturated" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002554_j.seppur.2020.117277-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002554_j.seppur.2020.117277-Figure2-1.png", "caption": "Fig. 2. Schematic diagram of (a) orifice membrane module, (b) cross-section view of membrane module, and (c) control, orifice plate with 24, 16, and 8 holes (top to bottom).", "texts": [ " A membrane module was designed to hold two flat membranes to let the porous support layer face the permeate side. Ultrafiltration membrane PAN 400 (Polyacrylonitrile, 200 kDa, ULTURA Co. Ltd., USA) with an effective filtration area of 24 cm2 (3 cm width, 8 cm length) was placed within the membrane module. On the frontal part of the membrane module, an orifice plate was designed and installed to generate turbulence via thrust power from the feed solution. The orifice plate was laser-punctured to have different numbers of holes with a diameter of 0.5 mm (Fig. 2). Orifice holes were fit into the flow channel (3 cm width, 0.2 cm height). Holes were aligned horizontally and spread evenly within the inlet area of the membrane for the sake of uniform turbulence generation throughout the membrane inlet. The numbers of holes used in this study were 24, 16, and 8. A plate with the effective opening area of 0.6 cm2, which was identical to that of the cross-flow inlet area, was used as control for membrane filtration. The total inlet areas of orifice plates were 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000608_icems.2019.8921477-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000608_icems.2019.8921477-Figure3-1.png", "caption": "Figure 3. Fault motor bearing system model", "texts": [ " When the transmission shaft system is twisted and deformed, the shaft system will generate a torque Ts, which is affected by the angle difference (\u03b81-\u03b82) and the rotating speed difference between the shaft motor end and the actuator end. When the bearing fails, the rolling body will create an impact when passing the faulty part, thus introducing a shock signal of specific fault frequency. The motor bearing fault can be equivalent to introducing an additional torque Tc [8]. The servo drive controls the motor to provide electromagnetic torque Te for the motor shaft. Te, Ts and Tc act on the motor shaft and affect the speed. The model of the fault motor bearing system is shown in Figure 3. The motor shaft inertia is J1 and the damping coefficient is b1. At the actuator end, the actuator has an equivalent moment of inertia J2 and a damping coefficient b2. The drive shaft torque Ts, the load torque TL and the additional torque Tc act together on the actuator to ultimately determine the speed of the actuator. According to the above theory establishes the following differential equations as relation (2) [9]: 1 1 1 1 2 2 2 2 1 2 1 2 1 1 2 2 / / ( ) ( ) / / e s c s L c s J d dt T b T T J d dt T b T T T c K d dt d dt (2) The system relationship (3) can be obtained by Laplace transform: 2 1 1 1 1 2 2 2 2 2 1 2 1 2 1 2 ( ) ( ) ( ) e c L c c J s T b s T T J s T b s T T T cs K T k t s s s s (3) The block diagram of the dual inertia transmission model as shown in Figure 4 can be derived by equation (3) for the establishment of the simulation model" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001997_gt2016-57905-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001997_gt2016-57905-Figure3-1.png", "caption": "FIGURE 3: CFD MODEL OF THE SMOOTH SEAL", "texts": [ " For each of the cases in Table 2, four exci- 4 Copyright \u00a9 2016 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/pdfaccess.ashx?url=/data/conferences/asmep/89521/ on 02/14/2017 Terms of Use: http://www.asme.org/about-asme/terms-of-use tation frequencies \u2126 are investigated. The used set of frequencies is detailed in Eqn. 16, in which \u2126rot = 166.67Hz corresponds to the rotational velocity of the shaft. \u2126 = [0.25,0.50,0.75,1.00] \u00b7\u2126rot (16) A full view of the implemented CFD model is presented in Fig. 3 which reflects one half part of two seals placed in a back-to-back configuration, sharing an inlet cavity. As seen in Fig. 3, the model includes the radial inlet cavity, the smooth seal and the outlet cavity. The calculations were performed using the commercial CFD TABLE 3: FLUID PROPERTIES Gaseous phase (air) Molecular Weight 28.96 [kgkmol\u22121] Dynamic Viscosity 1.831 [kgm\u22121s\u22121] \u00b710\u22125 Specific Heat 1.004 [J kg\u22121K\u22121] \u00b7103 Liquid phase (water) Density 997 [kgm\u22123] Dynamic Viscosity 8.899 [kgm\u22121s\u22121] \u00b710\u22124 Specific Heat 4.182 [J kg\u22121K\u22121] \u00b7103 solver ANSYS CFX. The fluids modelled in this study were atmospheric air for the gaseous phase and water for the liquid phase" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003719_s42417-021-00346-2-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003719_s42417-021-00346-2-Figure15-1.png", "caption": "Fig. 15 As mentioned on prior topics, the leading shoe is more required because the self-enforcement effect. In addition there is a tendency of this to be more unstable as can be seen on left side graph, that exhibits a more inclination than trailing shoe graph. However, FUF delta is low enough to justify great modifications on E shoe", "texts": [ " In addition it can be noticed that there are differences between the graphs exhibited on Fig.\u00a014. As can be seen, the graph 1 is indicates more unstability than 2 graph. This probably occurs because 1 refers to the friction coefficient of the shoe that suffers the self-enforcement effect. Despite this fact, the result obtained by these cases are similar, exhibiting the desired condition with Edrum values between 1.10 to 1.15E+11 Pa and 1 and 2 below 0.32 to obtain a TUF of 10.1. As can be seen on Fig.\u00a015, the values of FUF are very similar. However, the slope of these surfaces are different. The influence of 1 and 2 is weak and the FUF values are obtained owing to Eshoe variation. As noticed before, the leading shoe is more required because the self-enforcement effect. In addition there is a tendency of this to be 1 3 more unstable. The left side graph of Fig.\u00a015 that refer to Eshoe and 1 correlation confirms this exhibiting a more inclined surface than trailing shoe graph. Despite these differences, both situation exhibits the best condition with Eshoe between 1.73 and 1.83E+11 and all values of 1 and 2 above 0.35 to generate FUF of 2350 Hz. The worst condition is observed on regions that Eshoe is minimum and all values of 1 and 2 to produce FUF of 2317 Hz. As can be seen on Fig.\u00a016, the shoes must be stiff to not generate high eigenvalues. In addition 1 and 2 must be below 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002953_0954406220974052-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002953_0954406220974052-Figure6-1.png", "caption": "Figure 6. Illustration of passive gravity compensation on joint A based on the spring structure.", "texts": [ " However, due to the horizontal configuration of the device, the torque required on joint A relatively larger than those of the other two joints when conducting active gravity compensation. Since sG and rated torque smax are different on motors, the value of the overall output force is quite different. Also, it would be difficult to provide an adequate output force in the desired direction. Thus, the passive gravity compensation based on \u201cZero-free-length\u201d spring is adopted to reduce the torque consumption only for joint A as shown in Figure 6. Where Lx refers to spring elongation, r0 is the radius of the pulley, h0 is the mounting height of the pulley, LVS is the distance from the center of the pulley to the spring attachment point, LAS represents the distance from the active joint to the spring attachment point, and hi donates the angle between the active linkage and the horizontal direction, a0 is the angle between LAS and LVS, a1 represents the angle between LVS and Lx. From Figure 6, Lx can be calculated as Lx \u00bc \u00f0LVS 2 r0 2\u00de1=2 (17) According to the trigonometric function, LVS and a1 can be obtained by the formula LVS \u00bc \u00f0h02 \u00fe LAS 2 2h0LAScos\u00f090 hi\u00de\u00de1=2 (18) a1\u00bcarcsin r0 LVS \u00bcarcsin r0 \u00f0h02 \u00fe LAS 2 2h0LAScos\u00f090 hi\u00de\u00de1=2 (19) Similarly, a0 can be calculated as h0 sina0 \u00bc LVS sin\u00f090 hi\u00de (20) Therefore, a0 can be expressed as a0\u00bc arcsin h0sin\u00f090 hi\u00de \u00f0h02 \u00fe LAS 2 2h0LAScos\u00f090 hi\u00de\u00de1=2 ! (21) The torque generated by the spring on joint A can be calculated as equation (21) sS \u00bc FL \u00bc K1LxLASsin\u00f0a0 \u00fe a1\u00de (22) From equations (17) to (22), it can be seen that the torque generated by the passive compensation of the spring structure is related to the angle hi of joint A" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure4-1.png", "caption": "Fig. 4. Internal teeth type bearing.", "texts": [], "surrounding_texts": [ "The pitch and yaw bearings used in wind turbines have numerous bolt holes on their inner and outer ring, and they are respectively bolted to the hub and blade and the tower and nacelle during normal operation. To test these conditions, these bearings should be bolted to the corresponding frame. Furthermore, pitch and yaw bearings can be of internal or external teeth type, respectively shown in Figs. 4 and 5, depending on the location of the integrated gear teeth. Pitch bearings are typically of the internal teeth type, whereas yaw bearings can be of both types. Therefore, the test rig should be designed so as to be able to test both types of bearings. For internal teeth type pitch bearings, the outer ring is bolted to the hub and remains stationary, whereas the inner ring is bolted to the blade and rotates according to the wind speed. Therefore, the bearing\u2019s inner and outer rings should be divided into loaded and stationary parts to apply test loads to the loaded part while the stationary part is connected to the fixed supporting structure. The test rig should be able to load and rotate the test bearing simultaneously. It is impossible to perform the two functions simultaneously using only a set of bearings because of the structural configuration of the bearings and the test rig. To solve this problem, by using two sets of the same bearings, two rings with gear teeth attached for both bearings are con- nected to each other using an insert plate and rotated by the driving system whereas two rings without gear teeth attached are connected to the fixed and loaded frame of the test rig, respectively. In this case, between the two bearing sets, the bearing positioned lower is in an environment similar to the operational one. Figs. 6 and 7 show the connection methods for internal and external teeth type bearings, respectively. A bearing driving system comprising a driving motor, reduction gearbox, and pinion connected to the gearbox is used to rotate the test bearing in a manner following the actual environment. The driving systems for pitch and yaw bearings are generally called as pitch and yaw drives, respectively. The rotational speed of the bearing can be controlled by changing the gear ratio of the reduction gearbox and the number of pinion teeth." ] }, { "image_filename": "designv11_22_0001910_978-3-319-15010-9-Figure3.3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001910_978-3-319-15010-9-Figure3.3-1.png", "caption": "Fig. 3.3 (a) Network graph with external (control) agents r1 and r2 attached to agents v1 and v4 respectively, leading to an altered Laplacian A.G;R/ and input matrix B.R/ of model (3.2). (b) Equivalent electrical network. The potential difference Vv3 VR is the effective resistance between v3 and common resistor node fr1; r2g", "texts": [ "2, represent respectively, connection points between resistors corresponding to the edges E and ER. In addition, all connection points corresponding to the set R are electrically shorted. The effective resistance between two connection points in an electrical network is defined as the voltage drop between the two points, when a 1 Amp current source is connected across the two points. Then, the i -th diagonal element of A.G;R/ 1 is the effective resistance Eeff.vi / between the common shorted external agents R and vi . An example of the equivalent electrical network is displayed in Fig. 3.3. The implication is that J .G;R/ D 1 n nX iD1 Eeff .vi / : (3.5) Tree graphs are often adopted for agent-to-agent communication topologies as they minimize edge (communication) costs while maintaining connectivity. Using (3.5), we introduce some properties of J .G;R/ (3.4) specific to trees. 3For a survey of directed graphs, we refer the reader to [6]. 4All eigenvalues of L . QG/ are real, 1. QG/ D 0 and iC1. QG/ D i .A.G;R// for i D 1; : : : ; n. 44 3 Measures and Rewiring Let us first define the special set of agents that lie on any of the shortest paths between agents in R as the main path agents, designated by the set M: This is a unique set for a given pair ", "s/k2 as it is solely dependent on G and R. Similarly, we will denote P1.A.G;R/; B.R// as P.G;R/. Directly from the definition of the controllability gramian, one has kGG;R.s/k22 D tr \u02c6 1 0 eA.G;R/ B .R/ B .R/T eA.G;R/T d (5.12) D tr B .R/ B .R/T \u02c6 1 0 e2A.G;R/ d D 1 2 tr M .R/ A.G;R/ 1 ; (5.13) where M .R/ D B .R/ B .R/T . In the previous chapter a resistive electrical network interpretation was provided for the diagonal of the matrix A.G;R/ 1. An example of the equivalent electrical network is displayed in Fig. 3.3. We proceed to analyze this metric for our two special leader-agent cases; with jRj D jERj and jRj D 1 corresponding to Rd and Rc (defined in Sect. 2.2), respectively. Proposition 5.3. For a connected graph G, if each leader agent has exactly one edge and so each edge in Rd can have an independent signal then, kGG;Rd .s/k22 D 1 2 X vi2 .ER/ Eeff .vi / : (5.14) 5.3 Open Loop H2 Norm 93 Proof. This result follows directly from Lemma 3.13. ut Proposition 5.4. For a connected graph G, and all agents apply the same signal then kGG;Rc " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003483_j.jmapro.2021.03.054-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003483_j.jmapro.2021.03.054-Figure14-1.png", "caption": "Fig. 14. Schematic representation of the growth of the \u03b1 phase at the core area and the surface area in the EBM process: (a) Initiation of \u03b1-laths along with the \u03b1WGB at the core area, (b) No initiation of \u03b1-lath along with the \u03b1WGB near the surface area, (c) Basket-weaved \u03b1-laths structure at the core area, (d) Burger type \u03b1-colonies near the surface area.", "texts": [ " At first, \u03b1GB precipitated at the prior \u03b2/\u03b2 grain boundaries, and thereafter \u03b1WGB nucleated and developed into the \u03b2 grain to create a Widmansta\u0308tten \u03b1 colony during cooling, as shown in Fig. 13(a) and (b). The variant selection tendency in the core area was weaker due to the faster cooling rate and thus formed basket-woven \u03b1 laths, while the surface showed a strong variant selection with the help of a slower cooling rate, which led to the colonies and larger \u03b1 lath formations [16]. Therefore, some small \u03b1 laths appear in the \u03b2 grain in the core area, as shown in Fig. 13(a), and thus become larger and form a basket-woven structure that restricts the growth of \u03b1WGB, as shown in Fig. 14(c). On the other hand, due to the occurrence of strong variant selection during slow cooling, the basket-woven structure \u03b1 laths appear later and thus \u03b1WGB grows and forms a burger-orientation relationship colony, as shown in Fig. 14(d). Finally, it can be concluded that the \u03b1 transformation texture can be regulated with different cooling rates and thus the functionally graded part can be produced in the AM process. Fig. 15 shows the tensile properties of samples produced by the SLM and EBM processes. The results of the tensile test show that although both the Ultimate Tensile Strength (UTS) and the yield strength of the SLM samples are higher compared to the EBM samples, the modulus of elasticity (E) is opposite. There is a big difference between their UTS where SLM is above 1100 MPa and EBM is below 900 MPa" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure6-1.png", "caption": "Fig. 6. Equidistant-enveloping principle: (a) plane equidistant; (b) spatial equidistant.", "texts": [ " During the engagement motion, the contact point located on the conjugate curves moves along the line of action. The results show that these paired curves can satisfy the continuous motion and meshing conditions. The feasibility of the basic principle is further verified. 2.2 Tooth profiles modeling The conjugate curve theory shows that a pair of conjugate curves have continuous and tangent motion in the given contact direction. Based on equidistant-enveloping generation, the tooth surfaces can be obtained in terms of the designated conjugate curves. The equidistant-enveloping principle is shown in Fig. 6. General plane equidistant from curve T1 to curve T2 is displayed in Fig. 6(a). The curve T2 is obtained through the motion of curve T1 along the designated locus with distance L. For spatial equidistant in Fig. 6(b), the designated normal direction n in space curve trihedron is given and curve W1 is the original curve. Curves W2 and W3 are the corresponding equidistant curve of curve W1 with distance l1 and l2 in different normal direction, respectively. They can be expressed as 2 1 2 1 2 1 1 2 1 1 : , W W W W W W x x l W y y l z z l \u23a7 = + \u22c5 \u23aa = + \u22c5\u23a8 \u23aa = + \u22c5\u23a9 x y z n n n (21) and 3 1 3 1 3 1 2 3 2 2 : . W W W W W W x x l W y y l z z l \u23a7 = + \u22c5 \u23aa = + \u22c5\u23a8 \u23aa = + \u22c5\u23a9 x y z n n n (22) Generation of tooth profiles is further developed by sphere enveloping motion, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000349_ilt-03-2019-0088-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000349_ilt-03-2019-0088-Figure1-1.png", "caption": "Figure 1 Structural diagram of spiral groove liquid film seal", "texts": [ " So far, some studies have mentioned the effect of boundary slip or cavitation on liquid film seal, respectively. However, few investigations analyze the influence of boundary slip with cavitation on liquid film seal. In this study, a numerical model of liquid film seal based on Navier slip model and JFO boundary is established. The numerical model is used to investigate the effect of boundary slip on the hydrodynamic performance. To improve the load-carrying capacity, the optimized slip length is obtained. The structure of sealing face is shown in Figure 1(a). In the operating condition, the liquid from the inner groove radius ports moves toward the groove outer radius by the rotation of the seal rings until it is impeded at the edges of the land and the dam regions and elevated pressure is generated during the process (Lee and Kim, 2011). The rotating and stationary surfaces are separated when the opening force formed by the liquid film pressure exceeds or equals to the closing force. Between the seal rings [Figure 1(b)], the groove and land produce hydrodynamic pressure while the seal dam produces hydrostatic pressure (Basu, 1992). Where u L and uG represent the circumferential angles of the land and groove and u c is the angle of single-period computational domain. ri and ro represent the inner and outer radii. rgi and rgo represent the inner and outer groove radii. pi and po are the pressures at the inner radius and outer radius. a is the spiral angle. hg represents groove depth and h0 represents the film thickness on land" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002902_0954406220967693-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002902_0954406220967693-Figure2-1.png", "caption": "Figure 2. Tooth geometry of the virtual spur gear. (a) Stiffness arrangement in single pair of gear teeth in contact. (b) Stiffness arrangement in two pairs of gear teeth in contact.", "texts": [ " The latter parameter can be computed as Db \u00bc W n (1) Where: Db stands for the tooth width of the virtual spur gear, n denotes the number of the virtual spur gears. Finally, W denotes the tooth width of the straight bevel gear. Apart from the number of gear teeth and the tooth width, each virtual spur gear has unique tooth geometry parameters which can be written in function of the pitch angles of the straight bevel gear pair d1; d2\u00f0 \u00de. Then, for each virtual spur gear, the potential energy method has been used to compute the mesh stiffness. The concise steps to determine the mesh stiffness of the virtual spur gear can be described as follow: Figure 2 depicts the tooth geometry of the virtual spur gear. The expressions of shear stiffness, ks, of the gear tooth can be written as Ref.21 1 ks \u00bc 1:2 1\u00fe v\u00f0 \u00decos2 a1\u00f0 \u00de cos a2\u00f0 \u00de N 2:5 Ncos a0\u00f0 \u00de cos a3\u00f0 \u00de ! EDbsin a2\u00f0 \u00de \u00fe Z a2 a1 1:2 1\u00fe v\u00f0 \u00de a2 a\u00f0 \u00decos a\u00f0 \u00decos2 a1\u00f0 \u00de EDb sin a\u00f0 \u00de \u00fe a2 a\u00f0 \u00decosa da (2) The expressions of axial compressive stiffness, ka, of the gear tooth can be written as Ref.21 1 ka \u00bc sin2 a1\u00f0 \u00de cos a2\u00f0 \u00de N 2:5 Ncos a0\u00f0 \u00de cos a3\u00f0 \u00de ! EDbsin a2\u00f0 \u00de \u00fe Z a2 a1 a2 a\u00f0 \u00decos a\u00f0 \u00desin2 a1\u00f0 \u00de 2EDb sin a\u00f0 \u00de \u00fe a2 a\u00f0 \u00decos a\u00f0 \u00de da (3) The expressions of bending stiffness, kb, of the gear tooth can be written as Ref" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure3.17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure3.17-1.png", "caption": "Fig. 3.17 Projection onto a flat surface, with different projection systems", "texts": [ " However, this does not completely determine the three-dimensional orientation of the object, since the rotation about the forward direction is still unspecified. A third rotation is needed to completely determine the orientation of the object. This third rotation\u03c8 would not affect the direction in which the gun is pointing, it would only rotate around the pointing vector. For a quaternion solution to the targeting problem, see also Sect. 4.5.1. Another frequent paradigm is a projection onto a flat surface. Consider the following practical problem (Fig. 3.17). Two projection systems are mounted at a distance d in front of a flat surface, and should both project a point at P, located on the screen at the location (hor/ver), where the positive horizontal direction on the screen is to the right, and the positive vertical direction up. The lower system is mounted like an aerial gun: it can rotate about a vertical axis (\u03b8N ) and swivel about a (rotating) horizontal axis (\u03c6N ). The system is mounted below cm lower than the center of the screen-based coordinate system, and when \u03b8N = \u03c6N = 0 the projector is pointing straight ahead toward the screen", " 50 3 Rotation Matrices (The sign before hor is negative, because the positive direction for \u201chorizontal\u201d on the screen is to the right, but the corresponding positive direction for \u201chorizontal\u201d for the projector is to the left.) From that the corresponding nautical angles can be determined with Eq. (3.35) from the previous example with the aerial gun. For the upper projector, the target is at p = (d/ \u2212 hor/ver \u2212 above) b\u2032 x = p |p| and the projector angles can be found by applying the first two equations of Eq. (3.28): \u03b8H = arcsin ( py\u221a px 2+py 2+Pz 2 ) \u03c6H = \u2212 arcsin ( pz\u221a px 2+py 2+pz 2 \u00b7 1 cos \u03b8H ) . (3.36) A Python solution, with numbers approximating a setup such as Fig. 3.17, would be from skinematics import vector (d, hor, ver, above, below) = (1.5, 0.3, 0.2, 0.7, 1.4) p_lower = [d, -hor, ver+below] p_upper = [d, -hor, ver-above] lower_projector = vector.target2orient(p_lower, orient_type=\u2019 nautical\u2019) upper_projector = vector.target2orient(p_upper, orient_type=\u2019 Helmholtz\u2019) 3.6 Applications 51 Section 4.5.3 shows the solution of a somewhat more complex, but conceptually similar problem: orienting a camera in a missile such that it is directed on a selected target. An interpretation of the values of the rotation matrix can be found by looking at Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002953_0954406220974052-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002953_0954406220974052-Figure3-1.png", "caption": "Figure 3. The serial mechanism of HFFD-6 and its coordinate setup.", "texts": [ " According to the geometric relationship BiCij j2 \u00bc Lb 2, equation (3) is obtained to calculate the coordinate of the center point of the moving platform P\u00f0x; y; z\u00de R\u00fe Lasinhi r\u00f0 \u00decosgi x 2 \u00fe R\u00fe Lasinhi r\u00f0 \u00desingi y 2 \u00fe La2coshi z\u00f0 \u00de2 \u00bc L2 b (3) Three nonlinear equations from equation (3) with three unknown parameters can be obtained. Newton iterative numerical calculation equation (4) is used to solve the problem, according to the following equation, the coordinates of the moving platform can be obtained. Where xn represents the current approximate solution. xn\u00fe1 \u00bc xn f\u00f0xn\u00de f0\u00f0xn\u00de (4) As illustrated in Figure 3, a 3-DOF serial structure is developed to obtain a user\u2019s hand posture. Note that the three axes meet at the same point. Moreover, a counterweight is added to achieve a static balance of the rotating joint 4. Furthermore, the pen-style endeffector is designed to keep the rotating joint 6 maintains a static balance. The moving platform is fixed on the parallel mechanism, the angles of hi\u00f0i\u00bc4; 5; 6\u00de on three joints can be obtained by three encoders, respectively. Table 1 depicts the parameters of the serial structure, based on the D-H approach, the coordinate\u2019s frame of the linkages is established, and forward kinematics equations are obtained" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003057_iros45743.2020.9341633-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003057_iros45743.2020.9341633-Figure2-1.png", "caption": "Fig. 2. Diagram of an orbital robot (CAESAR [18], blue) on right. The forced dynamics of this system can be simulated on a fixed-base CAESAR manipulator (red) for HLS validation.", "texts": [ " And, \u03c42, \u03c42d \u2208 R n are the actuation and lumped torques (including gravity torques) of the manipulator, respectively. Also, C\u0304 \u2208 R (6+n)\u00d7(6+n) is the non-unique CC dynamic matrix and M\u0304 = [ M\u0304b M\u0304bq M\u0304\u22a4 bq M\u0304q ] is the coupled in- ertia1. J\u0304 \u2208 R 6\u00d7n, J\u0304b(q) \u2208 R 6\u00d76 are the end-effector Jacobians for the manipulator and vehicle, respectively, and Fe \u2208 R 6 \u223c= se(3)\u2217 is end-effector wrench. Def. 3: An orbital robot is a multibody system of n + 1 rigid links (including the spacecraft), which are connected with n holonomic-joints. The simplified configuration space of the orbital robot, as seen in Fig. 2 (blue box), is Q\u0302 \u223c= SE(3)\u00d7 R n with coordinates z = (gb, q) \u2208 Q\u0302. As in Sec. II-B, we also obtain a reduced Lagrangian for the orbital robot with the same system velocity 1M\u0304b, M\u0304bq, M\u0304q are the locked, coupling and manipulator inertias, respectively [7]. 1880 Authorized licensed use limited to: East Carolina University. Downloaded on June 17,2021 at 08:21:25 UTC from IEEE Xplore. Restrictions apply. V , but with an inertia tensor, M\u0302(q) \u2208 R (6+n)\u00d7(6+n), as l\u03022(q, V ) = 1 2V \u22a4M\u0302(q)V . The Hamel\u2019s equations describe its dynamics as, [ M\u0302b M\u0302bq M\u0302\u22a4 bq M\u0302q ] V\u0307 + C\u0302(q, V )V = [ F\u0302b \u03c4\u0302 ] + [ J\u0302b(q) \u22a4 J\u0302(q)\u22a4 ] Fe, (3) such that (3) has the same form as (2), however, without lumped torques/wrenches, and the corresponding terms are denoted with (\u2022\u0302) instead of (\u2022\u0304)", " 1 is capable of executing only the motion of an orbital robot\u2019s manipulator. So, the spacecraft pose, gb, is entirely reconstructed in software using the simulated actuation F\u0304b, and we denote this HLS configuration space as Q = SE(3)\u00d7 R n. Orbital robots can have a total arm length of 2 \u2212 4[m] [1], [18] depending on mission requirements. Firstly, the HLS validation of these unwieldy-sized robots while using a vehicle-manipulator poses additional complexities [4], which can be avoided using a fixed-base setup. This is illustrated in Fig. 2, where an OBSW is interfaced with the fixed-base CAESAR arm [18] (red box). Secondly, for HLS validation of sensor-based navigation, physical floating-base dynamics is also necessary. In this case, a scaled-down model might be used in a vehicle-manipulator. Therefore, in both cases outlined above, a common interface (Fig. 2, yellow) between an OBSW and OGFM is required, which maps the controller torques, (F\u0302b, \u03c4\u0302), to that of OGFM, such that the latter\u2019s statespace trajectories are identical to that of the orbital robot. Past HLS methods have relied on Hamel\u2019s equations for simulating the forced dynamics of the orbital robot. On one hand, this requires q\u0308 or \u03c4 measurements for reconstructing spacecraft motion (first row in (3)), which is a sensory overhead for the HLS. On the other hand, the Hamel\u2019s equations lack the momentum consistency of free-floating dynamics", " Taking the inertia-scaled difference in these accelerations, we obtain, \u039b\u0302q(q\u0308\u03b1 \u2212 q\u0308i) = \u2212EL(L, \u03be) + \u03c4\u0302 \u2212 A\u0302\u22a4 l F\u0302b + J\u0302\u22a4 g Fe \u2212 \u039b\u0302q\u2202q\u0307q\u0307(Li) \u22121 ( \u2212 EL(Li, X) + \u03c4i + \u03c4id + J\u22a4 i Fe ) (12) If \u03c4i is chosen as (11) and substituted in (12), we get, \u039b\u0302q(q)(q\u0308\u03b1 \u2212 q\u0308i) = 0n. In addition to this equality, if and only if, Ass. 1 holds true, we get identical trajectories for (q(t), q\u0307(t)), thereby proving equivalence of Li, L. To illustrate the result, for example, applying Theorem 1 on a fixed-base OGFM (i = 1) in (1), we get, \u03c41 =C(q\u0307)q\u0307 \u2212 \u03c4d \u2212M \u039b\u0302\u22121 q \u0393q(V )\u03be+ M \u039b\u0302\u22121 q ( \u03c4\u0302 \u2212 A\u0302\u22a4 l F\u0302b ) + ( M \u039b\u0302\u22121 q J\u0302\u22a4 g \u2212 J\u22a4 i ) Fe (13) This means that Theorem 1 enforces an interface between the OBSW and the OGFM such that the joint-space trajectories of both robots in Fig. 2 are identical. Theorem 1 also guarantees that the state-feedback (q, q\u0307) (red arrow) to the OBSW is consistent with expected orbital robot dynamics. 1882 Authorized licensed use limited to: East Carolina University. Downloaded on June 17,2021 at 08:21:25 UTC from IEEE Xplore. Restrictions apply. Theorem 2: (Converse matching for EP equations): Given locked dynamics of an orbital robot in (5) and vehiclemanipulator in (9), they produce the same equations of motion, i.e. (gb(t), \u00b5(t)), if and only if, Ass" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure14-1.png", "caption": "Figure 14. Voluminal Mesh of the disc and pads (39208 nodes, 20351 elements)", "texts": [ "The results of calculations of contact described in this section relate to displacements or the total deformation during the loading sequence, the field of equivalent Von Mises stress on the disc, the contact pressures of inner and outer pad at various moments of simulation.One proceeds then, the influence of some parameters on the computation results. The finite element model of the rotor is carried out with a mesh of 20351 elements for a total of 39208 nodes. The mesh of the disc and pad resulting from ANSYS software is presented in Figure 14. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. A test of convergence is designed to evaluate the influence of the mesh on the precision of the digital simulation. One tested four cases of mesh (fine, hexahedral and quadrilateral) whose characteristics are presented in Table 6. Figure 15 shows the meshing models of the couple disc-pads. According to Table 7, one notices that the equivalent maximum constraints of Von Mises increase, according to the number of elements of the grid" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003057_iros45743.2020.9341633-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003057_iros45743.2020.9341633-Figure1-1.png", "caption": "Fig. 1. On-ground facility manipulators. On left, a fixed-base manipulator. On right, a vehicle-manipulator with a mechanical mounting at the industrial manipulator end-effector, {B}.", "texts": [ " The problem statement of HLS for the orbital robot is described in Sec. III. In Sec. IV, the main idea is proposed using the method of Controlled Lagrangians (CL), and LP equations are introduced. In Sec. V, we propose two converse theorems for dynamics equivalence and three modalities for simulating spacecraft motion in HLS. In Sec. VI, we provide experimental results, followed by concluding remarks in Sec. VII. Def. 1: A fixed-base manipulator is a multibody with n holonomic-joints, see left of Fig. 1. Its configuration is denoted with coordinates, q \u2208 T n, corresponding to the joint positions and T n is a Riemannian manifold with manipulator inertia, M(q) \u2208 R n\u00d7n, as its metric tensor. Note that, for simplicity, the local isomorphism T n \u223c= R n, is used in control applications and also, henceforward in this paper. The Lagrangian for this manipulator is L1 = 1 2 q\u0307 \u22a4M(q)q\u0307 \u2212 U(q), where U(q) is the gravity potential. The dynamics are written using the Euler-Lagrange operator, EL(L1) [15, eq. 2.3] as, EL(L1, q\u0307) := M(q)q\u0308+C(q, q\u0307)q\u0307 = \u03c41+\u03c41d+J1(q) \u22a4Fe (1) where C \u2208 R n\u00d7n is the matrix of joint-space CC terms, and \u03c41, \u03c41d \u2208 R n are actuation and other lumped torques (including gravity potential torques, i.e. \u2202U(q) \u2202q ), respectively. J1(q) \u2208 R 6\u00d7n and Fe \u2208 R 6 \u223c= se(3)\u2217 are the manipulator Jacobian and wrench at end-effector, respectively. B. Vehicle-manipulator Def. 2: A vehicle-manipulator [16] (Fig. 1, right) is a multibody system consisting of a fixed-base manipulator of n holonomic-joints mounted on a vehicle at a frame {B}. The configuration of the manipulator and vehicle are denoted with coordinates, q, and gb \u2208 SE(3) (see App.), respectively. A vehicle might be another manipulator as shown in grey in Fig. 1. For model-based HLS, the vehicle-manipulator can be modeled as a fully-coupled system on the simplified configuration space, Q\u0304 \u223c= SE(3) \u00d7 R n. Considering the gravity potential forces as external, the presence of group coordinate gb yields a reduced Lagrangian [17, \u00a75] with inertia tensor, M\u0304(q) \u2208 R (6+n)\u00d7(6+n), as l\u03042(q, V ) = 1 2V \u22a4M\u0304(q)V , where V = [ V \u22a4 b q\u0307\u22a4 ]\u22a4 is the system velocity, V \u2227 b \u2208 se(3) (see App.) is the se(3) algebra for the vehicle pose, gb, and g\u0307b = gbV \u2227 b . Specifically, using l\u03042, the Hamel\u2019s equations describe the dynamics for the vehicle-manipulator [14, eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001922_acc.2016.7526819-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001922_acc.2016.7526819-Figure1-1.png", "caption": "Fig. 1. a) Geometric guidance principle for the 2D case; b) Geometric guidance principle for steady state situation", "texts": [ " For the desired value of the surge velocity we choose urd(ux, V\u0302) = \u221a (ux \u2212 V\u0302x)2 + (\u2212V\u0302y)2 (6) where V\u0302 is an estimate of the ocean current. Note that we assume that the ocean current is unknown and we use a linear kinematic observer, to be introduced in the next section, in order to get an estimate of the ocean current and of the global position of the vehicle. The steady state surge velocity urd,ss(ux,V) = \u221a (ux \u2212 Vx)2 + (\u2212Vy)2, which is reached when the vehicle is on the path, implies the desired along path velocity ux, see Fig. 1. The following assumption has to hold for the surge velocity Assumption 8: We assume |urd | \u2265 \u2016V\u2016 (7) Remark 5: This is a natural assumption in order to overcome the ocean current disturbance. To the end of guiding the vehicle towards the path we define a desired rotation matrix Rd b1d =ux \u2212 e\u0302\u2212 V\u0302 (8) Rd = [ b1d \u2016b1d\u2016 b2d \u2016b2d\u2016 ] = [ ux\u2212V\u0302x\u221a N ke\u0302+V\u0302y N\u0302 \u2212ke\u0302\u2212V\u0302y N\u0302 ux\u2212V\u0302x N\u0302 ] (9) where ux = [ux, 0]T is the desired constant along path velocity, e\u0302 = [0, ke\u0302]T gives the estimate of the cross-track error scaled by the constant gain k, V\u0302 = [V\u0302x, V\u0302y]T is the estimation of the current, and \u221a N\u0302 = \u2016V\u0302 + ux + e\u0302\u2016 = \u221a (ux \u2212 V\u0302x)2 + (\u2212ke\u2212 V\u0302y)2. The unit vector b2d \u2016b2d\u2016 is chosen to be perpendicular to b1d \u2016b1d\u2016 . The meaning of the vector b1d is clear from Figure 1. The vector b1d is the sum of e,ux,\u2212V\u0302. Therefore, no matter what is the sign or direction of the current, if the vehicle is not on the path it will have a non zero component pointing towards the path due to the vector e\u0302. In fact, when the estimate \u2212V\u0302 has converged, it cancels out the kinematic disturbance due to V. When e\u0302 is zero the final desired direction is such that the orientation of the vehicle counteracts the current and moves along the path with velocity ux. This situation is shown in Figure 1. To sum up: what we aim to do with our guidance approach is to define the direction b1d due to geometric considerations. This direction points towards the path and it is used to define a desired rotation matrix Rd which aligns the surge axis of the vehicle along b1d. In fact, since we consider under-actuated vehicles the only way for the vehicle to move with the assigned absolute velocity is to align its x axis along the direction of the desired absolute velocity. In this section we describe the observer used for the estimation of the ocean current" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000154_j.wear.2018.12.088-Figure18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000154_j.wear.2018.12.088-Figure18-1.png", "caption": "Fig. 18. Effects of running distance, abrasive media type, concentration compared to normal sliding and dry wear test (@ 5.37 Nm applied load).", "texts": [], "surrounding_texts": [ "The dry sliding gear wear test was performed as the worst scenario case baseline for worn surface and wear debris analysis and then compared to other wear modes. In general, the worn surface in comparison to an unworn surface had a severe adhesive wear pattern as shown in Fig. 24(a) and (b). In conjunction with that, severe adhesive wear particles, as shown in Fig. 25 were observed and confirmed that normal rolling and sliding occurred during the test performed under a similar condition to other tests. The wear debris was firstly extracted from its lubricating oil sample, this was achieved by filtration using membrane patch filter paper with 0.45 \u00b5m pore size. Fig. 20 shows a typical worn out surface of 10,000m running distance without lubricant supply. The worn surface exhibits failure wear modes such as scuffing and scoring. In conjunction to the worn surface, typical severe adhesive wear debris was observed as shown in Fig. 25. This is evidence that direct metal to metal contact was the major wear mechanism whereby the asperities of the two opposing gear surfaces were plastically deformed and weld together leading to transfer from one surface to the other." ] }, { "image_filename": "designv11_22_0002060_j.wear.2016.11.002-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002060_j.wear.2016.11.002-Figure1-1.png", "caption": "Fig. 1. Dry sand/rubber wheel abrasion apparatus (Adapted from ASTM G65).", "texts": [ " The two most commonly used abrasion tests are the jaw crusher gouging abrasion test (ASTM-G61) and the dry sand rubber wheel test (ASTM-G65) [9,10]. The jaw crusher gouging abrasion test is primarily used to study the wear of ground engaging tools interacting with hard and large abrasives representing conditions commonly associated with quarry and metallic mineral mining operations. For the case of fine abrasives such as oil sands, the dry sand rubber wheel (DSRW) test is more suitable as there is little occurrence of breakage during excavation in a soft abrasive medium like oil sands. The dry sand rubber wheel setup is shown in Fig. 1. The general procedure for the ASTM-G65 test consists of the following steps: cleaning and weighing the specimen; fixing the specimen in the Please cite this article as: Z. Lin, et al., Specific energy and the mod 10.1016/j.wear.2016.11.002i holder and loading a set force between the specimen and the rubber wheel; setting the revolution counter; adjusting and starting the sand flow; starting the wheel rotation; stopping the drive motor after running the desired number of wheel revolutions; and removing, cleaning and reweighing the specimen" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure7-1.png", "caption": "Fig. 7. Tooth profiles generation.", "texts": [ " 6(b), the designated normal direction n in space curve trihedron is given and curve W1 is the original curve. Curves W2 and W3 are the corresponding equidistant curve of curve W1 with distance l1 and l2 in different normal direction, respectively. They can be expressed as 2 1 2 1 2 1 1 2 1 1 : , W W W W W W x x l W y y l z z l \u23a7 = + \u22c5 \u23aa = + \u22c5\u23a8 \u23aa = + \u22c5\u23a9 x y z n n n (21) and 3 1 3 1 3 1 2 3 2 2 : . W W W W W W x x l W y y l z z l \u23a7 = + \u22c5 \u23aa = + \u22c5\u23a8 \u23aa = + \u22c5\u23a9 x y z n n n (22) Generation of tooth profiles is further developed by sphere enveloping motion, as shown in Fig. 7. The center of the selected sphere moves along the equidistant curve and the tooth profile can be obtained according to the enveloping method if given the parameter range. So an internal gear pair which has the single contact point between the mated gears can be developed. Equations of tooth profiles of internal gear pair are written as 2 2 2 2 2 2 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 1 cos cos cos sin sin ( , , ) ( ) 0 : { cos sin , cos cos ,0} { sin cos , sin sin , cos } { ' , ' , ' }, W W W W W W x x l y y l z z l t t l l l l l x y z t \u03d5 \u03b1 \u03d5 \u03b1 \u03d5 \u03d5 \u03b1 \u03b1 \u03d5 \u03d5 \u03b1 \u03d5 \u03b1 \u03b1 \u03d5 \u03b1 \u03d5 \u03b1 \u03d5 \u03d5 \u03a9 \u03a9 \u03a9 = +\u23a7 \u23aa = +\u23aa \u23aa = + \u23aa \u2202\u03a9 \u2202\u03a9 \u2202\u03a9\u23aa\u0394 = \u00d7 \u22c5 =\u23aa \u2202 \u2202 \u2202\u23aa\u03a9 \u23a8\u2202\u03a9 = \u2212\u23aa \u2202\u23aa \u23aa\u2202\u03a9 = \u2212 \u2212\u23aa \u2202\u23aa \u23aa\u2202\u03a9 =\u23aa \u2202\u23a9 (23) and 3 3 3 3 3 3 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 2 cos 'cos ' cos 'sin ' sin ' ( ', ', ') ( ) 0 ' ' ' : { cos 'sin ', cos 'cos ',0} ' { sin 'cos ', sin 'sin ', cos '} ' { ' , ' , ' } ' W W W W W W x x l y y l z z l t t l l l l l x y z t \u03d5 \u03b1 \u03d5 \u03b1 \u03d5 \u03d5 \u03b1 \u03b1 \u03d5 \u03d5 \u03b1 \u03d5 \u03b1 \u03b1 \u03d5 \u03b1 \u03d5 \u03b1 \u03d5 \u03d5 \u03a9 \u03a9 \u03a9 = +\u23a7 = + = + \u2202\u03a9 \u2202\u03a9 \u2202\u03a9\u0394 = \u00d7 \u22c5 = \u2202 \u2202 \u2202 \u03a9 \u23a8\u2202\u03a9 = \u2212 \u2202 \u2202\u03a9 = \u2212 \u2212 \u2202 \u2202\u03a9 = \u2202 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 (24) where ,\u03d5 \u03b1 and ', '\u03d5 \u03b1 are the sphere surface parameters" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001791_0954410016647076-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001791_0954410016647076-Figure1-1.png", "caption": "Figure 1. Three-dimensional interception geometry.", "texts": [ " Next, by using the Lyapunov function and backstepping method, threedimensional nonlinear guidance laws accounting for the autopilot lag are designed for several cases: (1) the target does not maneuver; (2) the target acceleration information can be acquired; and (3) the target acceleration is not known. Subsequently, simulation results are presented to show the applicability of the proposed methods. Some concluding remarks are made in the last section. Applications of planar guidance laws to three-dimensional guidance problem Formulation of a three-dimensional missile-target engagement Consider the spherical LOS coordinates \u00f0r, , \u00de with origin fixed at the missile\u2019s gravity center. Let \u00f0er, e , e \u00de be the unit vectors along the coordinate axes (Figure 1). By virtue of the principles of kinematics, the three relative acceleration components \u00f0ar, a , a \u00de can be expressed by the following set of second-order nonlinear differential equations16\u201318 \u20acr r _ 2 r _ 2 cos2 \u00bc aTr aMr ar \u00f01a\u00de r \u20ac cos \u00fe 2_r _ cos 2r _ _ sin \u00bc aT aM a \u00f01b\u00de r \u20ac \u00fe 2_r _ \u00fe r _ 2 sin cos \u00bc aT aM a \u00f01c\u00de In the process of terminal guidance, only the acceleration normal to missile\u2019s velocity can be adjusted and so we will just discuss the relative motion normal to the LOS. The purpose of designing a guidance law is to nullify the LOS angular rates _ and _ ", " By omitting the thirdorder terms in equations (27), (34), and (43), the NLG1 can be simplified as u1 \u00bc 2N_r2 \u00fe 1 r c1N_r x2 \u00fe 1 c1 \u00fe N_r r x5 \u00f045a\u00de u2 \u00bc 2N_r2 \u00fe 1 r c2N_r x4 \u00fe 1 c2 \u00fe N_r r x6 \u00f045b\u00de The NLG2 can be simplified as u1 \u00bc \" 2N_r2 \u00fe 1 r c1N_r x2 \u00fe 1 c1 \u00fe N_r r x5 \u00fe c1 N_r r aT \"3 sgn z3 # \u00f046a\u00de u2 \u00bc \" 2N_r2 \u00fe 1 r c2N_r x4 \u00fe 1 c2 \u00fe N_r r x6 \u00fe c2 N_r r aT \"4 sgn z4 # \u00f046b\u00de and the NLG3 can be simplified as u1 \u00bc 2N_r2 \u00fe 1 r c1N_r x2 \u00fe 1 c1 \u00fe N_r r x5 \u00fec1\"1sat \u00f0x2\u00de \u00fe N_r r \"1sat 0 \u00f0z3\u00de \u00f047a\u00de u2 \u00bc 2N_r2 \u00fe 1 r c2N_r x4 \u00fe 1 c2 \u00fe N_r r x6 \u00fec2\"2sat \u00f0x4\u00de \u00fe N_r r \"2sat 0 \u00f0z4\u00de \u00f047b\u00de We investigate a space interception problem. Define an inertial reference coordinate system which is parallel to the coordinate system, MXYZ, as shown in Figure 1. This system is inertially fixed and is centered at launch site at the instant of the launch. In this system, the X-axis is taken to be in the horizontal plane and in the direction of the launch, the positive Z-axis is in the vertical plane, and the Y-axis is chosen in such a way that the coordinate system forms a right-handed coordinate system. In this example, the initial target-to-missile range is r0\u00bc 15,000 m. The missile\u2019s initial position coordinates are xM0 \u00bc 0m, yM0 \u00bc 0m, and zM0 \u00bc 0m: Its initial velocity is VM0 \u00bc 1000m=s and its initial flight path and heading angles are \u2019M0 \u00bc 30 and M0 \u00bc 0 respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003557_s00501-021-01108-z-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003557_s00501-021-01108-z-Figure10-1.png", "caption": "Fig. 10: LMDStrategy (Zones1, 2&4)", "texts": [ " 9, noting that the approach for zone 4 is simply a mirrored copy of zone 1. For the LMD-p perspective, the left and right flash lands (zones 1 and 4)weredepositedwith3 layers, each at 0.7mm in height with a 1.2mm stepover between each bead. This created a slight excess in deposited material which would subsequently allow for a smooth blend between the addi- tive material and the base die when machined to the final specification. A similar approach was taken for the zone 2 flat cavity however only 2 deposit layers were required. Fig. 10 illustrates the strategy for these areas. Due to its more intricate geometry, the rear flash land required a more complicated approach for the deposition process. This involved having two stages, whereby the inclined surface was built first and then partially machined, before deposition on the top of the flash land. This approach resulted in an overall higher quality deposit, since it allowed for the deposition head to remain normal to the die substrate for the majority of the deposition. This was Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002946_j.sna.2020.112448-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002946_j.sna.2020.112448-Figure2-1.png", "caption": "Fig. 2. The lateral scanner a) CAD drawing b) manufactured device c) out- of-plane mode and d) torsion modes observed by Finite Element Analysis software, at 143 and 250 Hz respectively.", "texts": [ " NdFeB magnets were preferred as they are safe to use in human GI track applications [15] and are readily available in small sizes with low cost. The lateral scanner provides a combination of two orthogonal mechanical modes, thus offering Lissajous scan pattern, upon applying a weighted sum of both mode frequencies onto the coil. The scanning head dimensions as well as the tapered elliptical cross section of the flexure of the lateral scanner is tailored to set out-ofplane and torsional modes as the first two vibration modes, while pushing other undesired mechanical modes to higher frequencies and securing a good mode separation [16]. Fig. 2 illustrates the CAD drawing of the manufactured lateral scanner along with the first two mechanical modes (out-of plane and torsion). We refer the interested readers to other publications on the lateral scanner for a detailed description and characterization results [16,17]. Overall, the proposed device offers a hybrid \u2022 Length of scan head Lc 3 Flexure length Lf 4 Elliptical Flexure radii (tapered) a1,2; b1,2 0.5\u22120.25; 1\u22120. hydraulic and electromagnetic) actuation scheme in a compact orm factor (10 mm diameter) enabling laser steering in 3D" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure10-1.png", "caption": "Fig. 10. 3D view of test rig.", "texts": [ " The dimensions of the final test rig developed are 11 m \u00d7 10 m \u00d7 4 m (length \u00d7 width \u00d7 height), and Figs. 10 and 11, respectively show its 3D view and actual appearance. The main design parameter of the developed test rig was the safety factor with regard to the yield stress of the test rig material. The minimum value of 1.5 was selected as the target safety factor considering the large size of the test rig, and the design of the test rig was modified by trial-and-error to achieve this target. As a result of the modification, several support brackets were applied to test rig parts, as shown in Fig. 10. Structural analysis by using NX Nastran [10] was performed to check whether the design requirement of the test rig was satisfied. Target loadings were selected from extreme load cases that always show a much larger loading level than the fatigue load. Among the various extreme load cases listed in Table 1, the one having the least static load factor was used for the final evaluation loading. The static load factor is calculated as the ratio of the bearing\u2019s maximum allowed load to the external load applied to the bearing" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001472_s0263574714002458-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001472_s0263574714002458-Figure14-1.png", "caption": "Fig. 14. A few snapshots of a path of 160 configurations for grasping the teapot by a four-section continuum manipulator. (a) Initial Config. (b) Intermediate Config. 1 (c) Intermediate Config. 2. (d) Grasping Config.", "texts": [ "2 ms) is also much shorter than the time (18 ms) reported in,26 which conducts collision detection based on spheres and updates BVHs dynamically, for meshes with comparable numbers of triangles. We have also tested the performance of CD-CoM vs. OPCODE employed by a motion planner for grasping in a cluttered space by a continuum manipulator.47 The planner searches a path of configurations for an n-section continuum manipulator to gradually wrap around a target object with its tip following the contour of the object without colliding with nearby obstacles. Figure 14 shows snapshots of such a path, which consists of 160 configurations for a four-section continuum manipulator to grasp the teaport in a confined environment. To plan the path, the planner checked 1120 configurations. Figure 13 compares the collision detection time by CD-CoM vs. OPCODE for each collision check of the 1120 configurations of the four-section continuum manipulator in the search of the feasible path for grasping the teapot. Table V compares the average and median time per collision check and the total time of collision check needed for the 1120 configurations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000923_icpe.2015.7167962-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000923_icpe.2015.7167962-Figure7-1.png", "caption": "Fig. 7. Actual current distributions", "texts": [ " The position error of Theta-err smoothly reduces to zero in both cases by the proposed adaptive algorithm and also the back emf constant converges to its real value. The tested motor has no position sen d-axis current reference of 12.44A sh used in order to clearly interpret the q-a with respect to the given position error current distributions in the actual sync frame for the given position error. W position follows after the actual position 7(a), the torque producing current d current becomes negative so that th increases to cope with the same load a But in case of Fig. 7(b), the q-axis cu cope with the same load at the same spe positive torque producing current du current. However, when the estimated with the actual position, the d-axis curre the generated motor torque any more. When the positive position error de shown in the upper one of Fig. 6, t decreases to cope with the same load but the d-axis current maintains the sam same reference. However, the q-axis cu the lower one of Fig. 6 because the error decreases to zero and the positive current due to the d-axis current also red The algorithm is implemented on a TMS320F28335 floating point digital The switching frequency of the IGBT and the dead time is 1 . sor. Therefore the own in Fig. 3 is xis current change . Fig. 7 shows the hronous reference hen the estimated as shown in Fig. ue to the d-axis e q-axis current t the same speed. rrent decreases to ed because of the e to the d-axis position coincides nt does not affect creases to zero as he q-axis current at the same speed e value due to the rrent increases at negative position torque producing uces to zero. Texas Instrument signal processor. inverter is 10kHz control algorithm for s (Ke-init). The sampling time is 1ms 100 for both the current sensorless control" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003394_s10846-021-01352-y-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003394_s10846-021-01352-y-Figure5-1.png", "caption": "Fig. 5 CADmodeling view of the designed thrust test rig", "texts": [ " This moment is measured at the sting balance moment center and is transferred into the MAV aerodynamics center in the Simulink model by the following equation: Mac \u00bc Mb\u2212L d\u00f0 \u00de \u00f08\u00de Where Mac is the moment about MAV aerodynamics center is, Mb is the moment about the balance center, and d is the difference distance between balance and aerodynamic center. The moment coefficient changes with the elevator angle and angle of attack. The variation in the moment coefficient with flap deflection is linear but not constant from one flap to another one, as shown in Fig. 4(e). This variation will cause nonlinearity in the system response. A test apparatus is manufactured to measure the thrust of the MAV\u2019s propulsion system as a function of airspeed at various motor RPMs. A CAD modeling view of this test apparatus is shown in Fig. 5. An open jet wind tunnel with a maximum wind speed of 15 m/s is used with the test rig. Tests are performed on various motors and propellers to choose a set that is suitable for the fixed-wing MAV. This test provides useful information about the propulsion system before building the MAV. Applying the designed thrust test rig in the wind tunnel, the thrust for two RPMs is plotted versus speed from 5 m/s to 15 m/s, as shown in Fig. 6. As expected, the thrust decreases as the wind speed increases and increases with the propeller speed" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001850_amr.1137.61-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001850_amr.1137.61-Figure9-1.png", "caption": "Fig. 9 Effect of tool vibration in rotary USM: (a) Effect on hole quality, (b) Effect on rod/scrap formation", "texts": [ " A model parameter (which models the ratio of fractured volume to indented volume of single diamond particle) was shown to be invariant for most machining conditions. The accuracy of the model was demonstrated for magnesia stabilized zirconia. Shen et al. [70] developed the model for material removal rate using support vector fuzzy adaptive network (SVFAN) in RUM. Results also compared with that obtained by using fuzzy adaptive network and it had also been showed that the combined approach is a more effective algorithm for the modeling. Advanced Materials Research Vol. 1137 71 Fig. 9 shows the effect of tool vibration on the hole quality. Li.et al. [71] carried-out an experimental investigation of different process parameters (ultrasonic vibration, feed rate and spindle speed) in rotary ultrasonic machining of graphite/epoxy panel. Their conclusions were as follows; 1. Using ultrasonic vibration, quality of hole was improved and it made the more stabile operation. 2. Cutting force was significantly decreased when ultrasonic vibration used. Ya et al. [72] have presented mathematical model for the material removal rate" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000518_j.mechmachtheory.2019.103637-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000518_j.mechmachtheory.2019.103637-Figure3-1.png", "caption": "Fig. 3. A UPU-5-UPS parallel mechanism.", "texts": [ " Thus, the parallel mechanism with six limbs can be denoted as RbRaP1RaRb-5-UPS or UPU-5-UPS. The UPU limb connects the centers of the fixed base and the moving platform, and the five UPS limbs are arranged symmetrically around the UPU. In order to realize symmetric and non-redundant actuation arrangements, the five P joints in the five UPS limbs are selected as the actuation joints. The UPU limb provides the only constraint to the five-DoF parallel mechanism. The initial pose of mechanism is shown in Fig. 3. Because each UPS generates the six-DoF motion, its motion can be expressed by the whole set of finite screws [21] re- gardless of the dimensions, as{ S f,UPS } = { 2 tan \u03b8 2 ( s r \u00d7 s ) + t ( 0 s )\u2223\u2223\u2223\u2223\u03b8 \u2208 [0, 2\u03c0 ], t \u2208 R, s \u2208 R3 , |s| = 1, r \u2208 R3 } . (19) As a parallel mechanism\u2019 motion is the intersection of the motions of all its limbs, the motion of this parallel mechanism is obtained as{ S f,PM,1 } = { S f,UPU } \u2229 { S f,UPS } \u2229 \u00b7 \u00b7 \u00b7 \u2229 { S f,UPS }\ufe38 \ufe37\ufe37 \ufe38 5 = { S f,RbRaP1RaRb } (20) = { 2 tan \u03b8b,2 2 ( sb rB \u00d7 sb ) 2 tan \u03b8a,2 2 ( sa rB \u00d7 sa ) t1 ( 0 s1 ) 2 tan \u03b8a,1 2 ( sa rA \u00d7 sa ) 2 tan \u03b8b,1 2 ( sb rA \u00d7 sb )} , where rA is the position vector of the common rotation center of the first U joint (RbRa), and rB is the position vector of the common rotation center of the second U joint (RaRb)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001609_s10015-015-0223-z-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001609_s10015-015-0223-z-Figure1-1.png", "caption": "Fig. 1 a Four phases for the stance condition, DF, HS, FS, and DS. b Sequence of phases in bounding gait", "texts": [ " In this section, we explain the bounding gait and two robot models, which we used in this paper. During the bounding gait, left and right legs simultaneously kick the ground and roll and yaw movements of the body are relatively small due to the left and right symmetry of the leg movements. Therefore, we consider the left and right legs as one leg. Although the bounding gait is not necessarily observed as a preferred gait of animals, it is considered as a simple model of galloping gait [1]. The bounding gait has four types of phases for the stance condition, as shown in Fig. 1a. When the fore and hind legs are in the air, we named this phase as Double leg Flight (DF). When only the fore or hind leg is on the ground, we named these phases as Fore leg Stance (FS) or Hind leg Stance (HS), respectively. When both legs are on the ground, we named this phase as Double leg Stance (DS). Figure 1b shows the sequence of phases in a periodic bounding gait, which we investigate. This figure starts from an apex height in DF. The fore leg touches the ground, and then lifts off before the hind leg touches the ground. The robot returns to the apex height in DF after the hind leg lifts off the ground. Our simulation results did not obtain the bounding gait which has DS phase. In this paper, we used two physical models to investigate the role of flexibility of the body. One model consists of a rigid body and two massless springs (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002176_icra.2014.6907155-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002176_icra.2014.6907155-Figure7-1.png", "caption": "Fig. 7: (a) The occupancy test algorithm determines if a robot is inside a given triangle. If any angle between neighbors of u is greater than \u03c0, then u is outside of the triangle. (b) Diagram of triangle expansion controller regions between robots uL and uR, each with \u03c0 8 bearing resolution. A robot in region 1 rotates around uL or uR, so that it always ends in region 2. A robot in region 2 always moves in the direction where its two inner angles are decreasing. By doing so, the controller always guides a robot from a lower to an adjacent higher number region using only \u03b8L and \u03b8R. All sample trajectories (red lines) converge to the goal region.", "texts": [ " IMPLEMENTATION A high-level finite-state machine of triangulation construction is shown in Fig. 6. Two robots are initialized in the Frontier-Wall state and placed at the base-edge. All other robots begin behind the base edge in the Navigation state. Table I lists helper functions for all algorithms below. The navigation contains three states; Nav-Internal, Expand-Triangle, and Wall-Follow. A new robot, u, enters the network in the Nav-Internal state, and runs algorithm 1 to navigate to a frontier triangle. Line 2 runs an occupancy test function, shown in Fig. 7a, that returns the current triangle, Tc, that contains robot u, and its owner, o. If Tc is a nonfrontier triangle, then u moves to an adjacent triangle that is closer to (fewer hops from) the frontier (lines 10 to 11). Theorem 3.2 ensures that the owner of Tc is connected to owners of adjacent triangles, so u learns the hops of all adjacent triangles with a 2-hop message similar to the geometry message from Fig. 2b. If Tc is a frontier triangle (line 3) or null (only true if u has just crossed the base edge, line 6), then u will create a new triangle. The variables u.L and u.R are set to the left and right neighbors of the frontier edge (lines 4 and 7), and the robot changes its state to Expand-Triangle (lines 5 and 8). In the Expand-Triangle state, u runs algorithm 2. Line 2 computes the left and right inner angles to the frontier neighbors, \u03b8L and \u03b8R. Line 3 runs the triangle-expansion controller illustrated in Fig. 7b until u is in region 3. We lack the space for a complete description of the controller, so we sketch its operation here. When robot u enters region 3, if \u03b8L > \u03b8R, u first moves toward Bu(u.L)+\u03c0 until \u03b8R \u2265 \u03c0 3 . It then changes its heading toward Bu(u.R) + \u03c0, Algorithm 1 NAV-INTERNAL 1: while u.state =Navigate-Internal do 2: Tc \u2190 GETCURRENTTRIANGLE() 3: if ISFRONTIERTRIANGLE(Tc) then 4: (u.L, u.R)\u2190 GETFRONTIEREDGENBR(Tc) 5: u.state \u2190 Expand-Triangle 6: else if ISONLYBASEEDGE(N(u)) then 7: (u.L, u" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003123_msec2015-9240-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003123_msec2015-9240-Figure4-1.png", "caption": "FIGURE 4. THE MESH USED TO VERIFY THE HEAT INPUT MODEL, DISPLACEMENT BOUNDARY CONDITIONS (MARKERS ON THE \u2212x FACE, AND HEAT SOURCE SCANS (LINES ON THE +z FACE, NOT TO SCALE). THE HATCH SPACING BETWEEN SCANS IS 0.1 MM.", "texts": [ " (18) are the use of the elongated length c\u0303 in place of c, which stretches the distribution in the local z direction, and the additional 1 2 \u2206t term, which shifts the peak heat input from the end to the middle of the segment. An example of the three power distributions Q, Q, and Q\u0303 compared side by side in the local x-z plane is shown in Figure 3. In order to verify the accuracy and efficiency of the proposed heat input models, simulations are performed for the same process with Goldak\u2019s model and each of the new models with various time increment sizes. An 18\u00d718\u00d71.5 mm3 Ti-6Al-4V substrate is heated by five heat source passes, each proceeding in the +x direction (see Figure 4). The substrate\u2019s density is 4430 kg/m3, its Young\u2019s modulus is 200 GPa, its Poisson\u2019s ratio is 0.3, and its coefficient of thermal expansion is 9 \u00b5m/m\u00b7K. Temperature dependent conductivity, specific heat, and yield strength are shown in Figure 5. Outside the illustrated range of properties, the closest value is used. Complete stress relaxation is simulated at 690 \u25e6C [36]. 4 Copyright \u00a9 2015 by ASME Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 02/02/2016 Terms of Use: http://www" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001851_s12206-016-0614-7-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001851_s12206-016-0614-7-Figure2-1.png", "caption": "Fig. 2. Schematic diagram of the traditional taper leaf spring.", "texts": [ " 1 shows a schematic diagram of the taper leaf spring of tandem suspension when the taper leaf spring is installed. The intermediate and the rear axles are connected by the taper leaf spring, and both ends of the taper leaf spring are freely supported at the intermediate and rear axles. The middle of the taper leaf spring is connected to a frame through a U-bolt. The taper leaf spring of tandem suspension does not have a shackle, joints, and bushing elements compared with the traditional taper leaf spring. Fig. 2 shows a schematic diagram of the traditional taper leaf spring. The shackle of the traditional taper leaf spring allows the rotational and translational movement with respect to the frame when under the load. Therefore, exhibiting the rotational and translational motion in the traditional taper leaf spring model is important. However, the taper leaf spring of tandem suspension only permits the translational motion of the ends of the taper leaf spring. Thus, the contact and friction of two adjacent leaves at the ends are crucial for the development of the taper leaf spring model" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002328_j.mechmachtheory.2020.103873-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002328_j.mechmachtheory.2020.103873-Figure9-1.png", "caption": "Fig. 9. Mises stress on wheel flank.", "texts": [ " For saving the time cost, the nodes for calculated flank are different from those of un-working gear flank. This simulation only simulates the mesh behavior of drive side. The coast side is the same process. Based on the simulation of hypoid gear meshing, the results are shown in case1 and case2. The bending stress for pinion and wheel are calculated by quasi-static analysis method. The wheel flank was selected to show Mises stress and contact pressure results to save the post-processing. Case 1: Face-milled hypoid gear The Mises stress for wheel in two meshing moments are demonstrated in Fig. 9 . The maximum point is in the center point of contact ellipse. The 2D and 3D diagrams of contact stress are shown in Fig. 10 . The maximum value is in the center of gear flank. The bending stress on pinion and wheel flanks are depicted in Fig. 11 (a) and (b). The positive maximum value is in the root of gear flank. The negative maximum value is in the working part of gear flank. Those stress values represent the tension-compression stress in meshing process. The root stress results of pinion and wheel are shown in Figs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003324_j.mechmachtheory.2021.104261-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003324_j.mechmachtheory.2021.104261-Figure1-1.png", "caption": "Fig. 1. Structure of the 2-dof DTPM.", "texts": [ " Since the Hessian matrix only reflects the instantaneous performance of the mechanism, the method of the passing energy of virtual deformation within a finite range is proposed. In Section 4 , to fur- ther improve the kinematic stability, distribution of additional links is discussed. In Section 5 , the influence of positions and numbers of middle links on stiffness is also analyzed. A prototype is manufactured and its motion and stability experiments are performed in Section 6 . The proposed 2-dof DTPM is constructed by using the D-SLiMs, as shown in Fig. 1 . The middle link C 1 C 2 connects the two corresponding corner joints of the two identical parallel SLiM limbs, i.e. Limb 1 and Limb 2. It has been analyzed that, in the current assembly, the moving platform ( B 1 B 2 ) has two translational mobilities. Two rotary actuators are mounted at the joints on the fixed platform. Under their actuation, the moving platform can translate in the XY plane. First of all, kinematics of half of Limb 2, which is marked by the zigzag dashed segments on the right side of the mechanism, is considered. Origin of the coordinates is set at Point A 1 . In Fig. 1 , m indicates the number of zigzag segments between A 2 and C 2 , n is the number of the zigzag segments between A 2 and B 2 , 2 l 0 is the length of the SLiM link, l 1 indicates the length of connecting links B 1 B 2 and C 1 C 2 , \u03b82 is the angle between the driving arm and the vector A 2 B 2 , \u03b22 is the angle between the driving arm and the X-axis. Coordinates of joints B 2 and C 2 can be expressed by \u03b82 and \u03b22 , respectively, { X B2 = X _ B2 ( \u03b82 , \u03b22 ) = n l 0 cos \u03b82 cos ( \u03b22 + \u03b82 ) + l 1 Y B2 = Y _ B2 ( \u03b82 , \u03b22 ) = n l 0 cos \u03b82 sin ( \u03b22 + \u03b82 ) , (1) { X C2 = X _ C2 ( \u03b82 , \u03b22 ) = m l 0 cos \u03b82 cos ( \u03b22 + \u03b82 ) + l 0 sin \u03b82 sin ( \u03b22 + \u03b82 ) + l 1 Y C2 = Y _ C2 ( \u03b82 , \u03b22 ) = m l 0 cos \u03b82 sin ( \u03b22 + \u03b82 ) \u2212 l 0 sin \u03b82 cos ( \u03b22 + \u03b82 ) . (2) According to the definition of \u03b81 and \u03b21 in Fig. 1 , the coordinates of joints B 1 and C 1 are { X B1 = X _ B1 ( \u03b81 , \u03b21 ) = \u2212 n l 0 cos \u03b81 cos ( \u03b21 + \u03b81 ) Y B1 = Y _ B1 ( \u03b81 , \u03b21 ) = n l 0 cos \u03b81 sin ( \u03b21 + \u03b81 ) , (3) { X C1 = X _ C1 ( \u03b81 , \u03b21 ) = \u2212 m l 0 cos \u03b81 cos ( \u03b21 + \u03b81 ) + l 0 sin \u03b81 sin ( \u03b21 + \u03b81 ) Y C1 = Y _ C1 ( \u03b81 , \u03b21 ) = m l 0 cos \u03b81 sin ( \u03b21 + \u03b81 ) + l 0 sin \u03b81 cos ( \u03b21 + \u03b81 ) . (4) Geometric constraints between connecting links B 1 B 2 and C 1 C 2 are, { ( X B2 \u2212 X B1 ) 2 + ( Y B2 \u2212 Y B1 ) 2 = l 2 1 ( X C2 \u2212 X C1 ) 2 + ( Y C2 \u2212 Y C1 ) 2 = l 2 1 , (5) where X and Y indicate the coordinate components of the each joint", " (7) By substituting the locations of B 1 , B 2 , C 1 , C 2 , and (7) into (6) , and specifying \u02d9 \u03b21 and \u02d9 \u03b22 as input parameters, (6) can be described as J a \u23a1 \u23a2 \u23a2 \u23a2 \u23a3 \u02d9 \u03b82 \u02d9 \u03b81 \u02d9 \u03b22 \u02d9 \u03b21 \u23a4 \u23a5 \u23a5 \u23a5 \u23a6 = \u23a1 \u23a2 \u23a2 \u23a3 0 0 \u03c9 2 \u03c9 1 \u23a4 \u23a5 \u23a5 \u23a6 , (8) where J a = \u23a1 \u23a2 \u23a2 \u23a3 \u2212n sin \u03b21 \u2212n sin \u03b22 \u2212n ( sin \u03b21 + sin \u03b22 ) / 2 \u2212n ( sin \u03b21 + sin \u03b22 ) / 2 (1 \u2212 m ) sin \u03b21 \u2212(1 + m ) sin \u03b22 (1 \u2212 m )( sin \u03b21 \u2212 sin \u03b22 ) / 2 (1 \u2212 m )( sin \u03b21 \u2212 sin \u03b22 ) / 2 0 0 1 0 0 0 0 1 \u23a4 \u23a5 \u23a5 \u23a6 , is the Jacobian matrix. The determinant of J a is | J a | = 2 n sin \u03b21 sin \u03b22 . (9) When \u03b21 \u2208 { 0 , \u03c0} or \u03b22 \u2208 { 0 , \u03c0} , | J a | = 0 , the D-SLiM is singular, the kinematic performance is poor, which could lead to shakiness of the mechanism [34,40,41] . However, from the testing of the prototype, it is observed that idle motion appears not only in regions close to singularity. More causes of the idle motion are investigated in the following. The virtual deformations of connecting links B 1 B 2 and C 1 C 2 in Fig. 1 are defined as { L 1 = \u221a ( X B2 \u2212 X B1 ) 2 + ( Y B2 \u2212 Y B1 ) 2 \u2212 l 1 L 2 = \u221a ( X C2 \u2212 X C1 ) 2 + ( Y C2 \u2212 Y C1 ) 2 \u2212 l 1 (10) Then, the total virtual deformation energy of the two connecting links is E = E 1 + E 2 = 1 2 ( k 1 L 2 1 + k 2 L 2 2 ) (11) Let k 1 = k 2 = 1 , substituting the solutions of (5) into (11) , E reaches the minimum value. The minimization of (11) is consistent with (5) for the D-SLiM in Fig. 1 which only corresponds to partial workspace. In order to discuss the influence of multiple middle links on stability in the whole workspace, (5) is converted into the following energy minimization problem. The optimization model is formulated as Find : ( \u03b81 , \u03b82 ) min : E s.t. \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u23aa \u23aa \u23a9 X B1 \u2212 X _ B1 ( \u03b81 , \u03b21 ) = 0 , Y B1 \u2212 Y _ B1 ( \u03b81 , \u03b21 ) = 0 X B2 \u2212 X _ B2 ( \u03b82 , \u03b22 ) = 0 , Y B2 \u2212 Y _ B2 ( \u03b82 , \u03b22 ) = 0 X C1 \u2212 X _ C1 ( \u03b81 , \u03b21 ) = 0 , Y C1 \u2212 Y _ C1 ( \u03b81 , \u03b21 ) = 0 X C2 \u2212 X _ C2 ( \u03b82 , \u03b22 ) = 0 , Y C2 \u2212 Y _ C2 ( \u03b82 , \u03b22 ) = 0 ", " Since the initial value plays an important role in the Newton\u2013Raphson method, \u03b8ini , which corresponds to the workspace that the moving platform is only translating, is firstly adopted. Then, a perturbation is imposed on the initial value for searching other workspace components of the mechanism. It is known that a linkage could have multiple workspace components. In the following, relationship among the workspace components of the D-SLiM is discussed. With given \u03b21 \u2208 [1 \u25e6, 179 \u25e6] , \u03b22 \u2208 [1 \u25e6, 179 \u25e6] and \u03b21 + \u03b22 < \u03c0, for the 2-dof DTPM in Fig. 1 , three workspace components are obtained by solving \u03b8 with (13) in the domain [ 0 \u25e6, 9 0 \u25e6] . The initial values are given as \u03b81 = \u03b82 = ( \u03c0 \u2212 \u03b21 \u2212 \u03b22 ) / 2 . They correspond to the configurations that the moving platform is always parallel to the ground, i.e. \u03b3 = 0 . This workspace component is denoted by Q 1 in Fig. 2 (a). Thereafter, a perturbation is imposed on the initial value. According to the positive and negative values of \u03b3 , the solutions are divided into 2 categories. When \u03b3 < 0 , the moving platform is skewed to the right", " It is insufficient to explain the idle motion, only considering the angle deviations between each two configurations. To further estimate the stability of the mechanism, the investigation of the change of virtual deformation energy between different configurations is conducted in the following. In this section, an explanation for the instability of the mechanism based on deformation energy is presented. The eval- uation and optimization for the stability of the manipulator are performed by the energy function. Two example configurations of the 2-dof DTPM with one center-right middle link in Fig. 1 is discussed in this subsection. Example 1: \u03b21 = 1 . 0 rad and \u03b22 = 0 . 2 rad. The initial values, \u03b81 and \u03b82 , can be calculated from (7) . With a small perturbation imposed on the initial value, the virtual deformation energy of connecting links B 1 B 2 and C 1 C 2 around the prescribed configuration is calculated as shown in Fig. 7 . As can be seen from the detailed view of the bottom of the energy function E in Fig. 7 (a), there are three local minimums which are close. They correspond to three configurations in different workspace components, p 1 \u2208 Q 1 , p 2 \u2208 Q 2 , p 3 \u2208 Q 3 , respectively", " The maximum value of the path S is termed as the maximum passing energy, E m . E m = max { E( s i ) } (20) The length of S between p 1 and p 2 on \u03b81 \u2212 \u03b82 plane is L 12 = n \u2211 i =1 L ( s i ) (21) The average passing energy E is defined as E = ( n \u2211 i =1 E( s i ) \u00b7 L (s i ) ) / n \u2211 i =1 L (s i ) (22) The average passing energy reflects the performance of the idle motion within a finite range. For the same length of L (S) , the larger the average passing energy is, the smaller the idle motion is. Taking the mechanism in Fig. 1 as example, S 12 is denoted as the minimum cost path toward Q 2 from the ideal configuration, whose length on the \u03b81 \u2212 \u03b82 plane is L 12 . S 13 is denoted as the minimum cost path toward Q 3 , whose length is L 13 . Given L 12 = L 13 = 0 . 1 , the average passing energy and the maximum passing energy in the whole workspace toward Q 2 and Q 3 are presented in Figs. 12 \u201313 , respectively. Figs. 12 and 13 show that the average passing energy and the maximum passing energy have the same trend, though the values are different" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001535_s1068366615020063-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001535_s1068366615020063-Figure1-1.png", "caption": "Fig. 1. Schematic diagram of a mechanical face seal: (1) stationary ring; (2) rotating ring; (3) secondary seals.", "texts": [ " The importance of obtaining analytical expressions is also due to the fact that the use of the widespread computer programs based on the finite element method for calculating face seals is restricted because the hydrodynamic models they embody fail to take account of the specif ics of thin lubricant layers. The aim of this work it is to develop numerical and analytical techniques for calculating the hydrody namic characteristics of face seals with gaps of com plex form. Because of the deformation and nonuni form wear of the sealing rings, the gap in seals usually features tapering and undulation. NUMERICAL TECHNIQUE FOR CALCULATING THE HYDRODYNAMIC CHARACTERISTICS OF FACE SEALS When designing face seals (Fig. 1), one must con sider a wide range of problems, including the develop Keywords: face seal, gap, tapering, undulation, pressure distribution DOI: 10.3103/S1068366615020063 178 JOURNAL OF FRICTION AND WEAR Vol. 36 No. 2 2015 FALALEEV ment of a hydrodynamic pressure in the gap, friction in the gap, force and temperature load deformations of the friction pair, cooling of the friction pair, the rota tion of the shaft, and its axial displacements [2, 7, 8]. All of these problems are interrelated. The main task in designing a face seal for a turbomachine is to deter mine the hydrodynamic characteristics of the lubri cant layer in the gap" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002764_j.triboint.2020.106641-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002764_j.triboint.2020.106641-Figure9-1.png", "caption": "Fig. 9. Temperature change in the lubricant film and pads in the bearing midplane, n = 5000rpm, W = 2kN.", "texts": [ " The lubricant film thickness in Fig. 8 is not in scale, being 50 times magnified to see clearly. The minimum film thickness at the trailing edge of the bottom left pad was 36\u03bcm. The direction of rotation is considered clockwise, as indicated by the circumferential axis labels. A convergent lubricant film thickness was created at all the pads due to their tilting. Therefore, a sufficient supply of lubricant was ensured to prevent overheating. The temperature change in this bearing mid-plane is shown in Fig. 9. It was determined from conjugate heat transfer in the lubricant film and pads, described by energy equations (13) and (24). The boundary conditions mentioned in sect. 2.3 were applied. The temperature change of the assumed isothermal shaft was set to 20 \u25e6C, the heat flux at the leading and trailing free surfaces of the pads and at the outer surfaces of the pads was given by the heat transfer coefficient set to 115W/m2K. The mixing factor cmix = 0.65 was applied. The rectangular computational domain was discretised with the two uniform meshes with 21 nodes across the film thickness and 21 nodes across the pad thickness. The number of the nodes in the circumferential direction was the same as for calculating the pressure distribution. A mesh convergence study was conducted too. Fig. 9 is not in scale; the pads and film are modified to have the same thickness regardless of the actual lubricant film thickness. The maximum temperature change is in the location of the minimum lubricant film thickness, whilst the minimum is at the leading edges of the pads. The temperature above the leading edge of the pads changes parabolically towards the shaft; its mean value was obtained by mixing (sect. 2.4). In the lubricant film, the generated heat is transferred by convection and conduction" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000584_ffe.13150-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000584_ffe.13150-Figure4-1.png", "caption": "FIGURE 4 Schematic of positive and negative slip [Colour figure can be viewed at wileyonlinelibrary.com]", "texts": [ " This is done to avoid any error due to thermal expansion of test discs. The SRR ratio is defined as the ratio of sliding speed (Us) to rolling speed (Ur). The expression for calculating the SRR is given in Equation (1). It should be noted that, according to Equation (1), SRR is negative. Because of the negative SRR, the upper disc, which is slower, experiences negative slip, whereas the lower disc, which is faster, experiences positive slip. The schematic of the negative and positive slips is shown in Figure 4. It can be seen from Table 1 that the initial value of the lambda ratio (\u03bbin) defined as the ratio of the minimum film thickness (hmin) to the composite roughness (Sq \u00bc ffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi Sq21 \u00fe Sq22 q ) is 0.16 (boundary lubrication regime occurs for \u03bbin < 0.5) in this work, resulting in the running of the test discs under boundary lubrication regime. The expression for determining the minimum film thickness (hmin) for elliptical contact configuration under heavily loaded condition is developed by Hamrock33 (eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001174_pime_proc_1961_175_037_02-Figure24-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001174_pime_proc_1961_175_037_02-Figure24-1.png", "caption": "Fig. 24. Housing bore profiles", "texts": [ " 23 it will be seen that the reduction in bearing clearance at the butt faces was as much as the nominal clearance at the crowns, and it is for this reason that the bearing shells were eccentrically bored to give a larger clearance at the butt faces. These observations on the distortion of big-end bores are similar to those of Graham (11). Measurements of housing bore profiles at various stages of assembly revealed large bore distortions when the bolts were tightened. This is shown by comparison of X and Y in Fig. 24u and was caused by the offset position of the bolts in the butt faces. In Fig. 25u showing the bore profiles of the bearing shells when inserted in the rod, the distortions were considerable, producing ovalities of between 0.007 and 0.008 in. When reading these curves it should be appreciated Vol17.5 No I0 1961 at UNIV OF PITTSBURGH on August 14, 2015pme.sagepub.comDownloaded from 51 8 E. A. BLOUNT Static loading fixture E Double-acting loading cylinder. F Support plates. G Cap. H Probe. assernbIy with cap loading Proc Iwtti Mech Engrs that they only show deviations from the tolerance band and that, because of the large scale used, the distortions in shape are exaggerated", " The strength of the cap arch was increased to offset the increase in the length of arch benveen the points of fixing. The stability of the whole assembly was improved by locating the bolts at the centre of area of the butt faces. To test this design a model was flame-cut from mild-steel plate, the dimensions being to scale except that the width was half that of the engine rod and there was a single bolt each side instead of two bolts. Apart from the bore and butt faces, which were ground, the surfaces were left as cut. The distortion on tightening the bolts without the bearing shells is shown in Fig. 24b, and on tightening them with the bearing shells fitted in Fig. 25b. In both instances the distortion was much less than with the original assembly. Even more significant was the fact that when the maximum load was applied the ovality was less than half that with the original assembly, the increase in bore diameter being 0.0037 in., and the reduction across the butt faces 0,0032 in. For a split big-end assembly of 74 in. diameter this was considered satisfactory. Given these facts, the engine designers produced a design for the complete big-end assembly and this was manufactured, and tested for distortion in the same way" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002350_tpwrd.2020.2980163-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002350_tpwrd.2020.2980163-Figure10-1.png", "caption": "Fig. 10. Insulation from the ground of a shield wire: (a) schematic diagram; (b) picture of the layout in practice.", "texts": [ " The landform along the line was found to be complicated and changeable. The sags in the conductor and shield wire varied significantly across large spans between 87#-88# and 89#-90#. Therefore, the phase conductor and the shield wire had significantly different heights above the ground along the line used for capacitive power tapping. 2) Field Measurement Procedure: Initially, one shield wire was grounded at each tower, and the other shield wire was segmented and insulated from the ground via a post insulator in parallel with the protective air gap, as shown in Fig. 10. A single grounded point was set at tower 86# using a shortcircuit wire. In order to ensure the security of the test equipment and appliance during the field measurement of the equivalent circuit parameters, a ground rod was connected to the shield wire, and the original short-circuit wire at tower 86# was unconnected. During the measurement, the ambient temperature was 15 \u2103, the wind speed was 0~0.2 m/s and the average load current was 160 A. Authorized licensed use limited to: UNIVERSITY OF BIRMINGHAM" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure11-1.png", "caption": "Figure 11. Contact area", "texts": [], "surrounding_texts": [ "The numerical simulations using the ANSYS finite element software package were performed in this study for a simplified version of a disc brake system which consists of the two main components contributing to squeal the disc and the pads. Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. a) Boundary conditions applied to the disk In this FE model, the disk is rigidly constrained at the bolt holes. The bolt holes are tied as rigid body to a reference point, where the rotation of the disk is allowed in the y-axis, its angular velocity is imposed and constant \u03c9= 157,89 rad/s. The support is applied to the hole of the disk and is of cylindrical type of which the degrees of freedom are shown in Matrix 1. b) Boundary conditions and loading applied to the pads Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. The boundary conditions applied to the pads are defined according to the movements authorized by the caliper. Indeed, one of the roles of the caliper is to retain the pads which have the tendency natural to follow the movement of the disk when the two structures are in contact. The caliper maintains also the plates in direction Z. Thus, the conditions imposed on the pads are: \u2022 The pad is embedded on its edges in the orthogonal plan on the contac surface, thus authorizing a rigid movement of the body in the normal direction with the contact such as one can find it in an automobile assembly of brake (Coudeyras;2009) \u2022 A fixed support in the finger pad. \u2022 A pressure P of 1 MPa applied to the piston pad. \u2022 The contact between the pads and the disk has a coefficient of friction is equal to 0.2 . Boundary conditions in embedded configurations are imposed on the models (disc-pad) as shown in Fig. 12 (a) for applying pressure on one side of the pad and Fig.12 (b) for applying pressure on both sides of the pad. Matrix 1. 1 Radial Free Axial Fixed Tangential Fixed Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited." ] }, { "image_filename": "designv11_22_0002627_0954406220945728-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002627_0954406220945728-Figure5-1.png", "caption": "Figure 5. Analysis of the joint interface between A- and B-axis.", "texts": [ " By analyzing the six motion axes of the AFP machine, it can be seen that except for X- and Caxes, the motion of the remaining four axes will cause gravity deformation. Therefore, based on the model proposed above, corresponding joint interface is replaced by an elastic beam, and the schematic diagram is shown in Figure 4, where the red parts represent the abstract of elastic beam deformation models. Analysis of the joint interface between A- and B-axis According to kinematic chain, the deformation of the joint interface between A- and B-axis is affected only by the motion of A-axis. Thus it can be simplified to a twodimensional situation, as shown in Figure 5(a), where the blue line represents elastic beam deformation model and its length is LB, the elastic modulus is EB, the crosssectional area is AB, and the moment of inertia is IB, the circle H represents the fiber placement head of weight Gh. Moreover, the fiber placement head and the elastic beam are connected by a rigid beam of length Lh, and the motion of A-axis is described by . The analysis of the rigid beam and the elastic beam is shown in Figure 5(b) and (c) respectively, and the force and moment at the free end of the elastic beam can be obtained GB \u00bc Gh \u00f08\u00de MB \u00bc M \u00bc GhLh sin \u00f09\u00de Therefore, the displacement at the free end of the elastic beam is zB \u00bc Gh EBAB LB \u00f010\u00de yB \u00bc MB LB \u00fe zB\u00f0 \u00de 2 2EBIB \u00bc GhLh LB \u00fe zB\u00f0 \u00de 2 sin 2EBIB \u00f011\u00de XB \u00bc MB LB \u00fe zB\u00f0 \u00de EBIB \u00bc GhLh LB \u00fe zB\u00f0 \u00de sin EBIB \u00f012\u00de According to the elastic beam deformation model, the gravity deformation matrix between Aand B-axis is B ATg\u00bc 1 0 0 0 0 c XB s XB yB 0 s XB c XB zB 0 0 0 1 2 6664 3 7775 \u00f013\u00de Evidently, coefficients in the above model such as lengths, modulus, and cross-sectional area of the elastic beam are unknown as well as difficult to obtain" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003606_j.ijpvp.2021.104434-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003606_j.ijpvp.2021.104434-Figure7-1.png", "caption": "Fig. 7. System of reference adopted to represent the scanner measurements.", "texts": [ " The work piece topography captured with the 3D scanner is shown in Fig. 6; about 18 million points were acquired to describe the vessel geometry and successively elaborated by means of the ad hoc codes written in Matlab, confirming a satisfactory accuracy of the manufacturing process. In order to be coherent with the optimization procedure proposed in Section 2, in the following measurement results are presented by considering different sections of the vessel, identified by the azimuth angle \u03b8\u2208[0,\u03c0] describing a plane passing through the x-axis of the shell (see Fig. 7). For each section the position of a generic point of the profile is defined by means of the x coordinate or considering the angle \u03b3\u2208[0,\u03c0] defined between the x-axis and the line passing through O\u2032 lying on the plane defined by \u03b8 (see Fig. 7). Fig. 8(a) shows the measured points superimposed to the theoretical profiles (internal and external) for \u03b8 = 0 and \u03b8 = \u03c0, while Fig. 8(b) reports the corresponding normalized error er,m = em/L measured at the internal and external surfaces, where em is the difference between theoretical and real profiles. Results are quite satisfactory: only close to the ends an error of 0.4% of the axial length L is observed, whereas the remaining portion presents an error less than 0.2%. Similar results are obtained in the case \u03b8 = \u03c0/2 (see Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003075_icpei49860.2020.9431433-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003075_icpei49860.2020.9431433-Figure3-1.png", "caption": "Fig. 3. Motor test bench implemented in this work", "texts": [ " The estimated torque can be scaled to actual torque value by performing a calibration process. To evaluate the generated regression model, two error assessment methods are used; Root Mean Square Error (RMSE) and R value (Regression). R value is the linear correlation coefficient [9] as described in eq.1. between actual value and estimated value from the model where 1.0 is the perfect estimation case without error. = \u2211 (\u2211 )(\u2211 )[ \u2211 (\u2211 ) ][ \u2211 (\u2211 ) ] (1) III. EXPERIMENTAL SETUP To perform experiments of the proposed method, a motor test bench depicted in Fig.3 is constructed to use in both regression model generation and estimation performance evaluation processes. The actual value of rotor torque is obtained by a force sensor installed directly to the motor stator as illustrated in Fig.3 with the calculation of eq.2. = (2) The motor output is coupled with a speed reducer gearbox with the ratio of 100:1. The final stage output is at the gearbox shaft where a position encoder is coupled to observe output shaft speed and position. The encoder is used only for estimation performance evaluation and model validation. The motor is driven by a motor driver board DRV8305 [10] with a Texas Instrument microcontroller TMS320F28069 [11]. MATLAB Simulink is used to graphically programming the signal processing and machine learning sections of the proposed system to implement the complete FOC architecture and the torque estimation method", " The motor was controlled to run at three constant speeds: 700, 1000, and 1200 rpm to observe the prediction performance at different motor speed. Estimated torque values were obtained by four prediction models including Electromagnetic Torque and three regression methods: Neural network regression, Linear regression, and Stepwise regression. The results of torque sensor and regression model are compared for performance evaluation validation. 138 Authorized licensed use limited to: UNIVERSITY OF WESTERN ONTARIO. Downloaded on May 26,2021 at 11:59:13 UTC from IEEE Xplore. Restrictions apply. The motor test bench as shown in Fig 3. is attached with a motor under the test and target loads. Fig. 4 illustrates the procedure of regression model generation in the experiments. The electrical signals of the motor and sensor data, i.e. force, shaft speed and position are recorded in the Feature collection block. The regression model training process uses the collected data to generate prediction modes that are then can be used to estimate torque. The data set is divided into 2 sets for training and testing procedures with the ratio of 80:20" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001997_gt2016-57905-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001997_gt2016-57905-Figure2-1.png", "caption": "FIGURE 2: HALF PART OF SEAL HOUSING ASSEMBLY SHOWING INLET AND OUTLET FEATURES", "texts": [ " This is done by measuring the peak-to-peak time delay \u2206T between the harmonic position and force data series resulting from the CFD simulations. This can then be turned into a phase lag using \u03c6 = 2\u03c0\u2126\u2206T . The amplitudes of the forces fz and fy can be quantified as the mean of the maximum values of the peaks in a force data series. The seal geometry used in the numerical investigations presented in this paper is based on a simplified representation of the smooth annular test seals found in the multiphase seal test facility presented in [34]. A picture of the test facility vertically split seal housing assembly design can be seen in Fig. 2. As evident from Fig. 2 the seals in the test facility are mounted in a back-to-back configuration to alleviate axial thrust. This setup is adopted in the CFD analysis presented here. Additionally, the operating conditions imposed in the CFD analysis reflect a subset of the operating conditions under which the physical test facility can be used to identify rotordynamic coefficients for the test seals. Table 1 lists the geometrical dimensions of the seal as well as the operating parameters. The operating conditions listed in Table 1 are the same for all simulations presented in this paper" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure13-1.png", "caption": "Fig. 13. Loading spots for analysis.", "texts": [ " The final analysis model contained 321127 mesh elements\u2014124970 shell elements for the top frame main and side assembly; 152476 solid elements for the bearings, adaptor, and insert plate; 43175 shell elements for the base frame assembly; and 506 rigid elements for the boltconnected positions\u2014which have a hexagonal or tetragonal shape according to the location, and the mesh size was in the range of 50-60 mm. All hydraulic cylinder loadings were applied to the corresponding positions of the test rig, as shown in Fig. 13. From the analysis results, the maximum equivalent stress of the test rig was 167.3 and 171.2 MPa for a 3.0 MW pitch bearing and 2.5 MW yaw bearing, respectively, as shown in Figs. 14 and 15. This occurred at the bracket area of the top frame main assembly for both bearings. Because all parts of the test rig were composed of SM490A with thickness in the range of 16-40 mm and a yield strength of 315 MPa, the safety factors are 1.88 and 1.75 for the pitch and yaw bearings, respectively. Therefore, the test rig satisfies the design requirement and is structurally safe for each bearing\u2019s static loading conditions as well as fatigue loading conditions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002506_012078-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002506_012078-Figure1-1.png", "caption": "Figure 1. The shape of the contact patch of an elastic toroidal tire with a rigid supporting surface", "texts": [ " That is, in the case when the velocity vector of the wheel axis makes an angle , called the slip angle, with the plane of rotation of the wheel, namely: relationships for finding the magnitude of the lateral force, the stabilizing moment and the friction-loss power in the contact area are derived. To solve this problem, we will use the methodology and equations obtained earlier [1\u20137] for a wheel with a cylindrical treadmill (the shape of the contact patch is close to rectangular). DS ART 2019 IOP Conf. Series: Materials Science and Engineering 832 (2020) 012078 IOP Publishing doi:10.1088/1757-899X/832/1/012078 In order to simplify, we assume that the shape of the contact patch is close to elliptical (Figure 1), i.e., the coordinates of the contour line of the contact patch are related by: + = 1, 1 where A and b are the semi-axes of the ellipse, and the normal pressures in both the longitudinal and transverse sections of the contact patch are distributed according to a parabolic law [8]: , = 2 . 2 As in the above studies, we believe that there are two zones in the contact patch \u2013 an adhesion section at the contact entry and a sliding section at the contact exit. In the adhesion section, the longitudinal and transverse tangential displacements of the point of the wheel treadmill, due to the action of the longitudinal and transverse tangential forces, respectively, are defined as [5]: = \u2206 = \u2212 , 3 = \u0394 = = \u2212 \" \u2248 \u2212 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003719_s42417-021-00346-2-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003719_s42417-021-00346-2-Figure2-1.png", "caption": "Fig. 2 Drum components and coordinates system\u2014five parts: the drum, including the central part, and rotating at , two shoes in arcs, and two frictional linings localized at the exterior radius of the shoes", "texts": [ " This provides the best input parameters values that generates low values of unstability. The input parameters are one brake component Young\u2019s modulus and one frictional coefficient. The output parameters are FUF, the first unstable frequency, MUV, the maximum unstability eigenvalue and TUF, the total number of unstability frequencies. Despite the fact that automotive brake system have internal shoes, this arrange is a variation of long external shoe brake [26, 28]. Hence, the same of its design procedure can be applied to automotive drum brake (Fig.\u00a02), an infinitesimal circular section which will be used to estimate the contact pressure distribution. Pires [5] explained that applying forces directly on lining arc medium point results in appreciable error. The arc angle accounted for it is the difference between 1 and 2 . According to Norton [26], the normal pressure p at the contact is proportional to the distance between the shoe abutment point and its application. The normal pressure p can be expressed by [5, 26, 28]: where a is a proportionality constant and an angle defined in Fig.\u00a02. The origin of is along the axis which starts at the center of the drum and crosses the pivot of the shoe. The shoe lining goes from 1 to 2 . max is between and corresponds to the angle value where the contact pressure is maximum that can be expressed as (1)p = a sin( ), Relating Eqs. (1) and (2) leads to the profile of pressure, p( ) along the contact, where Eq.\u00a03 expresses the pressure profile over the lining area, thus max is the angle where p is max. Norton [26] suggest that max is the minimum angle value between 2 and 90\u25e6 . For the geometry considered (Fig.\u00a02), 1 = 10\u25e6 and 2 = 100\u25e6 , therefore, max = 90\u25e6 . pmax is the maximum admissible pressure of the lining material. Hence, the Eq. (3) is only subordinated to . However, this modification would not warrant some increase in performance that justifies the mass and inertia increase [26]. Assuming unidirectional sliding and permanent contact, the Coulomb\u2019s law of friction writes: where ft is the frictional force, fn is the normal force and is the friction coefficient. According to Norton and Orthwein [26, 28] fn is expressed by the following equation: where w is the lining width" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000741_978-3-319-75277-8-Figure6.2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000741_978-3-319-75277-8-Figure6.2-1.png", "caption": "Fig. 6.2 Recordingmovements of the lower armwith opticalmarkers. (Left)pi indicate the position of the markers, \u201c\u25e6\u201d the middle of the markers, and \u201c*\u201d the location of the Center of Mass (COM). (Right) Active markers for 3-D position measurements, for the Optotrak-system", "texts": [ " And the estimation of dynamic 3-D pose based on optical motion capture systems is described in (Selbie and Brown 2018). The presentation here focusses on the kinematic principles underlying the 3-D analysis. To define position and orientation of an object in three dimensions, one needs to find the positions of three points pi(t) that are firmly connected to the object. The only requirements for these points are that they (a) are visible, and (b) must not be arranged along a line. In the example sketched out in Fig. 6.2, threemarkers are attached to the left lower arm. Assume that the positions of these markers have been recorded, and stored as \u00a9 Springer International Publishing AG, part of Springer Nature 2018 T. Haslwanter, 3D Kinematics, https://doi.org/10.1007/978-3-319-75277-8_6 85 pi(t), i = 0, 1, 2. To investigate the object dynamics, the following questions have to be answered: 1. What are the positions x(t), linear velocities v(t) = dx(t) dt , and linear accelerations acc(t) = d2x(t) dt2 of the markers, with respect to our chosen space-fixed coordinate system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000672_robio49542.2019.8961446-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000672_robio49542.2019.8961446-Figure6-1.png", "caption": "Fig. 6\uff0eThe simulation model in ADAMS", "texts": [ " In the simulation, we only change the coordinates of X. Through PSO algorithm, we get f that maximize the adjusted cosine similarity. Simulations show that DMP can reasonably generalize any trajectory and converge to the target point. In order to verify the modified DMP method for trajectory planning and force control of the arm, we performed the MATLAB-ADAMS co-simulations. The simulation model was created in ADAMS, and the control algorithm was implemented in MATLAB and Simulink. The initial state of the simulation model is shown in Fig. 6. The simulation system contains a 7-DOF robotic arm, a 6-D force sensor, a tool, and an object. The arm would operate object on the curved surface of the object. In the simulation, we measure the force and torque of the fixed pair as the data of the six-axis force sensor. The simulations include DMP learning from demonstration in the given plane and DMP trajectory reproducing in an unknown surface. The demonstration data is given directly in MATLAB. The trajectory \u201c2\u201d is in the plane shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure7.11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure7.11-1.png", "caption": "Figure 7.11 Right-hand and left-hand lathe tools", "texts": [ " The cutting angles of many cutting tools are established during their manufacture and cannot be changed by the user. D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 7 Cutting tools and cutting fluids 107 7 same direction. Front clearance is also required to clear the workpiece surface. A tool used to part-off or form undercuts requires rake and clearance in the direction of feed but also requires side clearance to prevent rubbing in the groove produced, Fig. 7.10(b). the left, Fig. 7.11. a true rake angle somewhere between the two, which is the angle along which the chip will flow when cutting in either direction. The trail angle is required to prevent the rear or trailing edge of the tool from dragging on the workpiece surface. Cutting tools which are used to cut in only one feed direction require only one rake angle, although a number of clearance angles may be required to prevent rubbing. The knife tool shown in Fig. 7.10(a) acts in the direction shown, and a rake and a clearance angle are required in the True rake Trail angle Side rake Side clearance Back rake Front clearance Figure 7" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003847_j.mset.2021.08.005-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003847_j.mset.2021.08.005-Figure3-1.png", "caption": "Fig. 3. (a) Equivalent Principal Elastic Strain in workpiece during extrusion", "texts": [ " Therefore, ample lubrication of the die channel will result in a substantial reduction in the fluctuation of extrusion load and helps in efficiently reducing the amount of energy consumed throughout the actual extrusion process, resulting in better mould existence. For the current analysis, Fig. 2 shows that for a die having corner angle W = 20 and channel angle U = 90, the maximum value achieved is about 10.007kN for the process. (b) Maximum Principal Elastic Strain vs Time curve during simulation. Fig. 3 (a) explains the Equivalent Maximum Principal Elastic strain distribution in the workpiece during the extrusion process and 3 (b) illustrates the Equivalent Maximum Principal Elastic strain vs Time curve during simulation in the workpiece during extrusion. The zone in which the highest value of elastic strain is occurring inside the corner region of the channel is known as the main deformation zone. It can be detected from Fig. 3 (b) that the highest value for effective elastic strain which is achieved during the process varies from almost 0.003 to 0.3037. Fig. 4 (a) demonstrates the Effective Equivalent Total deformation distribution in the workpiece during the extrusion process and 4 (b) demonstrates the Equivalent Total deformation vs Time curve during simulation in the workpiece during extrusion. The zone in which the total deformation occurs inside the corner region of the channel is known as the main deformation zone" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002798_s00170-020-06152-6-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002798_s00170-020-06152-6-Figure10-1.png", "caption": "Fig 10 Combination of rectangular (B1,2) and circular (C1) datum target areas for the establishment of datum B and C", "texts": [], "surrounding_texts": [ "The well-known \u201cmaximum utilization\u201d rule concerning datum hierarchy states that \u201cIf a Datum Feature can and may constrain a degree of freedom, it must.\u201dWhen a DF is used to establish a datum, it constrains, or locks, some degrees of freedom of an ideal feature (e.g., the tolerance zone). As per ISO 5459, the maximum number of degrees of freedom that can be constrained by this integral feature is equal to, or less than, six minus the invariance degree of the nominal integral DF. However, high level of flexibility is again here provided to the designer, concerning the designation of the number of locked or released degrees of freedom in each datum of a datum system. This can be specified by the assignment of complementary indication ([PL], [SL], [PT], ><, [ ] or [Tx,Ty,Tz,Rx,Ry,Rz]) added after the datum identifier symbol in the relevant datum section, as shown in the example given in Fig. 7." ] }, { "image_filename": "designv11_22_0001239_0954410014537240-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001239_0954410014537240-Figure2-1.png", "caption": "Figure 2. Inertial frame and body frame.", "texts": [ " The wireless communication modem is used to monitor the status of the helicopter. In this work, a hybrid model is used to describe the dynamics of a small-scale helicopter, which includes the nonlinear rigid body dynamics, main rotor dynamics, and simplified yaw dynamics. The model uncertainties and other trivial factors are treated as external disturbances. The position of the gravity center of the helicopter is notated by P\u00bc [x, y, z]T in the inertial frame, where the linear velocity of the gravity center in the inertial frame is given by v\u00bc [vx, vy, vz]T (see Figure 2). For the orientation, we take advantage of roll( ), pitch( ), and yaw( ) representation defining \u00bc [\u2019, ]T and \u00bc \u00bd , , T. The angular velocity of the helicopter represented in the body frame is ! \u00bc \u00bd!1, r T, where !1 \u00bc \u00bd p, q T. Actually, using the Euler angles to represent the attitude kinematics, there are singularities when the pitch angle equals to 90 . In this paper, it is assumed that and are far away from 90 . at University of Sydney on March 13, 2015pig.sagepub.comDownloaded from The nonlinear rigid body dynamics of the fuselage can be derived by Newton\u2013Euler equation and written as _P \u00bc v m _v \u00bc R\u00f0 \u00deFb \u00femge3 _ \u00bc \u00f0 \u00de" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000709_psc49016.2019.9081455-Figure18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000709_psc49016.2019.9081455-Figure18-1.png", "caption": "Fig. 18. linear induction motor which is used for Maglevs [8].", "texts": [ " Fig. 16 demonstrates this method [13]. Recently, as high temperature superconductors are becoming more and more commercialized, the superconductive materials like YBCO and BSCCO are being tested as the material for the windings. This will simplify the cooling system. Fig. 17 compares the cooling system of a low temperature superconductive coil and a high temperature one. B. Linear Induction motors The type of linear induction motor which is used for Maglevs is short armature. As demonstrated in Fig. 18, LIMs are composed of on board armature windings and a conductive layer installed on the track and as the electrical energy needs to be transferred into the train, the electric connectors are necessary, limiting the speed of train. On the other hand as air gap of maglev trains, compared to rotational motors, is usually higher, the power factor and efficiency of these motors is normally very low [8]. Another difference between the radial and linear induction motors, is the presence of flux leakage at the edges of motor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001875_978-4-431-54595-8-Figure2.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001875_978-4-431-54595-8-Figure2.6-1.png", "caption": "Fig. 2.6 Two-joint model with three antagonistic pairs of PAMs. PAMs 1 and 2, PAMs 3 and 4, and PAMs 5 and 6 are paired, respectively. For simplicity, it is assumed that the moment arm of each joint is constant and that the characteristics of PAMs are the same. This model replicates an antagonistic structure with multiple muscles in a human upper limb", "texts": [ " The A-A concept is similar 2 Motor Control Based on the Muscle Synergy Hypothesis 35 to EPH in focusing on EP control; however, when we use the A-A ratio R and AA sum S in the agonist-antagonist system, the EP is represented linearly using the A-A ratio R and the joint stiffness G is also represented linearly using the A-A sum S (see (2.20) and (2.25)). These parameters can be controlled individually for a single-joint agonist-antagonist system. Consider a two-joint PAM model that mimics a human arm structure with shoulder and elbow joints and three antagonistic pairs of muscles around and connecting the two joints. The structure of the PAM model and its parameters are illustrated in Fig. 2.6. We assume the initial state at which the shoulder-joint angle is s D 0, the elbow-joint angle is e D 0, and the internal pressure and length of PAM are Pi D P0 and li D L .i D 1; 2; : : : ; 6/, respectively. When s and e are at the equilibrium state, the contraction forces Fi of PAMs i balance each other, such that F1 F2 C F3 F4 D 0 (2.26) F3 F4 C F5 F6 D 0: (2.27) 36 H. Hirai et al. Considering (2.8), (2.26) and (2.27) can be rewritten as K.P1/.l1 l0.P1// K.P2/.l2 l0.P2// C K.P3/.l3 l0.P3// K", " Find a couple of examples of the Bernstein problem of degrees of freedom in everyday movement. The motor redundancy problem may exist not only in a joint space but also in a muscle space. 2. The EPH has been a controversial theory of motor control for about half a century. Survey this remarkable theory and summarize its pros and cons. 3. Verify Eq. (2.8) from the input-output relationship of energy in a single PAM. 4. Considering the two-joint arm model with three antagonistic pairs of PAMs in Fig. 2.6, derive Eq. (2.39) which describes the relationship among EPs, A-A ratios, and A-A sums. 5. Consider the role of one motor command, the A-A sum, and describe what kinds of tasks its control is effective for. 6. Motor synergies may be described in terms of three features: sharing, flexibility/stability (error compensation), and task dependence (Latash 2008b). PCA is a powerful technique for detecting the first feature of synergies. What methods are available for testing the second feature of synergies" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002953_0954406220974052-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002953_0954406220974052-Figure10-1.png", "caption": "Figure 10. Illustration of the three-coordinate force measuring platform.", "texts": [ " Then, with the proposed active and passive combined gravity compensation approach, the end effector of the device is manually pulled to these positions randomly. The experimental results show that HFFD-6 can stay stationary and balanced at these positions with any configuration. Moreover, the effectiveness of the proposed gravity compensation approach is verified quantitatively by measuring the gravity compensation errors of the device at different positions. For this purpose, a three-coordinate force measuring system is constructed. As illustrated in Figure 10, the system is composed of four parts: (1) an ASUS laptop, (2) an intelligent display instrument MCK-F, (3) a three-dimensional force sensor, and (4) the gantry type three-axis motion platform. The control signals are transmitted from the ASUS laptop to the control system through USB communication, the data of the force sensor is displayed and transmitted to the laptop. For the convenience of the experiments, the serial end-effector is tentatively removed and its gravity is subtracted from the active gravity compensation model in equation (10)", " And this is the advantage gained due to the approximately equivalent torque available in three joints by adding the spring on joint A. As illustrated in Figure 13, the Fcon1 is larger for E2 than that for E1, this indicates the combined gravity compensation approach improves the maximum continuous output force of HFFD-6 at the sampling point. Secondly, the maximum continuous output force in the whole workspace,Fcon2, is calculated. Based on the scatter points obtained in the whole workspace of HFFD-6 (Figure 10), points are traversed to calculate the maximum continuous output force one by one. Among all the values, the minimum is recorded as Fcon2. The results are shown in Table 4. From Table 4, in E1, when the active gravity compensation approach is adopted, the maximum continuous output force Fcon2 in the whole workspace of HFFD-6 is 3.67N. While in E2, it increases by 54.8%, to the value of 5.68N with the proposed combined gravity compensation approach. It proves that the maximum output force capability of HFFD-6 is greatly improved by adding passive compensation to balance the output toque capacity of all three motors" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001519_s12046-015-0427-x-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001519_s12046-015-0427-x-Figure2-1.png", "caption": "Figure 2. General approach.", "texts": [ " This is physically and geometrically justified, since under load, the raceways shift toward each other and thus simultaneously altering the distance between their centers of curvature and compressing the rolling elements figure 1. By using this approach, it is possible to determine not only the equivalent contact load, but also the variation in the contact angle from the displacement of the curvature centers. A rolling element in a sector is defined by two nodes representing the centers of curvature of the raceways. Each node is linked to the corresponding opposite node by nonlinear traction springs or connector element (figure 2). The two zones of contact between the rolling element and raceway in a sector are modeled by rigid shells and coupled by rigid beam elements to the corresponding centers of curvature at two nodes materializing the contact ellipse, as shown in figure 2. It should be noted that local deformations, which are taken into account by the nonlinear connector elements or springs, should not be perturbed by further singular or numerical deformations. The rigid shell elements are added in order to minimize this risk, especially at the linkage zone of the rigid beams. The modeling of rolling elements by means of a connector element is a challenging task. The contact characteristics of the rolling element and the raceways have to be investigated carefully" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000518_j.mechmachtheory.2019.103637-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000518_j.mechmachtheory.2019.103637-Figure4-1.png", "caption": "Fig. 4. A 5-UPU parallel mechanism.", "texts": [ " (19) As a parallel mechanism\u2019 motion is the intersection of the motions of all its limbs, the motion of this parallel mechanism is obtained as{ S f,PM,1 } = { S f,UPU } \u2229 { S f,UPS } \u2229 \u00b7 \u00b7 \u00b7 \u2229 { S f,UPS }\ufe38 \ufe37\ufe37 \ufe38 5 = { S f,RbRaP1RaRb } (20) = { 2 tan \u03b8b,2 2 ( sb rB \u00d7 sb ) 2 tan \u03b8a,2 2 ( sa rB \u00d7 sa ) t1 ( 0 s1 ) 2 tan \u03b8a,1 2 ( sa rA \u00d7 sa ) 2 tan \u03b8b,1 2 ( sb rA \u00d7 sb )} , where rA is the position vector of the common rotation center of the first U joint (RbRa), and rB is the position vector of the common rotation center of the second U joint (RaRb). Thus, the mechanism has the same 3T2R motion with its UPU (RbRaP1RaRb) limb. 4.2. Example B If five identical SKCs that generate the second-kind 3T2R motion are used as limbs, a new parallel mechanism with the same 3T2R motion will be synthesized. For instance, a parallel mechanism composed by five RbRaP1RaRb limbs is shown in Fig. 4. The five limbs are arranged symmetrically. We denote each limb as UPU by supposing that the adjacent Ra and Rb joints share the same rotation center. In this way, the parallel mechanism is denoted as 5-R RaP RaR or 5-UPU. b 1 b At the initial pose of the parallel mechanism shown in Fig. 4, it should be noted that all the Ra joints in the five limbs of the mechanism have the same directions, so do all the Rb joints. The five limbs provide identical constraints to the moving platform. Hence, the mechanism is an over-constrained one. The symmetric and non-redundant actuation arrangements for the mechanism are realized by selecting the five P joints as the actuation joints. The motion of the parallel mechanism is the same as that of each RbRaP1RaRb, as{ S f,PM,2 } = { S f,RbRaP1RaRb } \u2229 \u00b7 \u00b7 \u00b7 \u2229 { S f,RbRaP1RaRb }\ufe38 \ufe37\ufe37 \ufe38 5 ,= { S f,RbRaP1RaRb } (21) which is equivalent with Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001442_20140824-6-za-1003.02165-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001442_20140824-6-za-1003.02165-Figure2-1.png", "caption": "Fig. 2. Two-link arm manipulator", "texts": [ " Under weak assumptions (see Van der Schaft [2000]) the system (18) is passive with supply rate \u03b8\u0307\u2032u and it is zero-state detectable with respect to [q, \u03b8]. In particular the matrix entries mi,j(q) of M(q) and ci,j(q, q\u0307) of C(q, q\u0307) are m1,1(q) = a1 + 2a3 cos(q2) + 2a4 sin(q2) m1,2(q) = m2,1(q) = a2 + a3 cos(q2) + a4 sin(q2) m2,2(q) = a2, c2,2(q, q\u0307) = 0 c1,1(q, q\u0307) = \u2212h(q)q\u03072, c2,1(q, q\u0307) = h(q)q\u03071 c1,2(q, q\u0307) = \u2212h(q)(q\u03072 + q\u03071) where h(q) = a3 sin(q2)\u2212a4 cos(q2), a1 = I1 +m1l 2 c1 +Ie + mel 2 ce + mel 2 1, a2 = Ie + mel 2 ce, a3 = mel1lce cos(\u03b4e) and a4 = mel1lce sin(\u03b4e) (see Fig. 2). The model parameters for the nominal model are m1 = 1, l1 = 1, me = 2, \u03b4e = \u03c0/6, I1 = 0.12, lc1 = 0.5, Ie = 0.25, lce = 0.6, J1 = 0.6, J2 = 0.5, k1 = 1, k2 = 1.5, \u03b21 = 0.01 and \u03b22 = 0.05 in the appropriate units of measure. Introducing the state variable x , [\u03b8, \u03b8\u0307, q, q\u0307] the state and control constraint sets are X ={x \u2208 R 8 | \u03b8i, qi \u2208 [\u22125\u03c0/6, 5\u03c0/6], \u03b8\u0307i, q\u0307i \u2208 [\u221250, 50], i = 1, 2}, U ={u \u2208 R 2 | ui \u2208 [\u22121000, 1000], i = 1, 2}, The robot is initially at rest with x0 = 0 and it is required to reach the steady state q\u03041 = \u03b8\u03041 = \u03c0/3 and q\u03042 = \u03b8\u03042 = \u03c0/2 while satisfying the constraints" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000439_j.mechmachtheory.2019.103612-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000439_j.mechmachtheory.2019.103612-Figure5-1.png", "caption": "Fig. 5. Bench press model.", "texts": [ " [55,56] while it has been modified in the following points: (i) the rotational inertia of the upper limbs segments is included, and (ii) the net joint moment and forces can be computed. In addition, a more meaningful expression to calculate the power is proposed. The bench press exercise can be realized with a traditional bench press or in a Smith machine. The mathematical model described is based on a Smith machine [57] where the barbell is guided and restricted to remain horizontal at any time. The upper limbs segments are considered as rigid bodies and joined through ideal joints, see Fig. 5 . The bench press exercise is modeled as follows. The arm of length L a is joined with revolute joints to the forearm and trunk, additionally the vertical displacement at the shoulder is allowed. The forearm of length L f is articulated to the hand through a revolute joint. The hand can be modeled as a point of lumped mass with the same displacement as the barbell. Since a Smith machine is used, two prismatic joints at each side of the barbell are included to allow the vertical displacement. Thus, the hand is only able to move vertically and the horizontal distance between the shoulder and the arm, d , remain constant. Also, since the barbell is forced to keep centered and horizontally, the bench press exercise is assumed to be performed in a symmetrical manner [56] . Based on this assumption, any half of the model depicted in Fig. 5 can be used to analyse the dynamics of both halves of the symmetrical model during the weight lift. In Fig. 5 , y is the vertical position of the hand, y s is the vertical position of the shoulder and \u03b8 is the angle between the forearm and the arm. The model described has two d.o.f, this means that knowing a set of two independent coordinates one can determine the mechanism position. Usually, in a bench press exercise the hand position (coordinate y ) is measured using optical encoders [58] or linear transducer [59] and the elbow angle coordinate \u03b8 is measured using a goniometer [55] . An inverse kinematic analysis is required to derive the position, velocity and acceleration of the limb segments from the measured time curves of the hand height, y , and elbow angle \u03b8 . In Fig. 5 , the arm and forearm define a triangle with vertices A, B and C . From the cosine theorem the distance AC can be written as: AC 2 = d 2 + (y \u2212 y s ) 2 = L 2 a + L 2 f \u2212 2 L a L f cos \u03b8, (8) where L a and L f are the limb segment lengths, d is the horizontal distance between wrist and shoulder and y s is the position of the shoulder axis. Using the previous equation, the shoulder axis position, y s , can be computed from the constants of the model and the values of y and \u03b8 . In addition, if the shoulder axis displacement can be assumed to be negligible, Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002599_s0263574720000685-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002599_s0263574720000685-Figure7-1.png", "caption": "Fig. 7. Planar ballbot system.", "texts": [ "1 m, g = 9.8 m/s2, and T = 15 s. The initial state is (2, 0, 0, 0) and the initial values of the control parameters d\u0302i (t) are chosen to be zero. Under the constructed controller, the system outputs x1, x2, x3, and x4 converge to zero, as illustrated in Fig. 6. The solid curves stand for the trajectories of the displacement x1 and velocity x2 of the cart, whereas the dashed curves for the angular displacement x3 and angular velocity x4 of the pendulum. The planar ballbot system,14 illustrated in Fig. 7, is a mobile robot that moves on a rolling hoop. The generalized coordinates of the system are denoted by the hoop angle \u03c6(t) and the body angle with respect to the ground \u03b8(t), both of which are increasing counter-clockwise. The red dot on the hoop is its initial contact point to the ground. One revolution of the initial contact point implies a 2\u03c0R linear displacement for the center of the hoop. Note that at the starting configuration (\u03c6(0)= 0), the https://www.cambridge.org/core/terms. https://doi" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001655_tmag.2014.2332431-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001655_tmag.2014.2332431-Figure1-1.png", "caption": "Fig. 1. Experimental device.", "texts": [ " A great advantage is that this kind of power supply offers total freedom in control of frequency and amplitude on each spectral component at any time. Secondly, the experimental results are exposed. It has been shown that according to the results obtained, the injection of HF signal over LF in the same magnetic circuit influence the LF iron loss in a way that they decreases. Finally, an interpretation of this complex phenomenon is discussed and a conclusion is drawn at the end of this paper. The experimental device is presented in Fig. 1. The toroidal magnetic circuit of 27 mm high, has been assembled with electrical steel sheets using nonoriented (NO) quality steel of 0.5 mm thickness, referenced in this paper as NO50. It has a primary winding of n = 100 turns uniformly distributed that is supplied by a voltage vp constituted by the superposition of two sinusoidal signals (vp = vp L + vp H ): 1) a LF signal vp L at fL frequency which, unless otherwise stated, is 50 Hz; 0018-9464 \u00a9 2014 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002350_tpwrd.2020.2980163-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002350_tpwrd.2020.2980163-Figure4-1.png", "caption": "Fig. 4. Schematic diagram of the horizontal distance calculation process.", "texts": [ " Thus, the arithmetic average values are taken as the equivalent circuit parameters of capacitive power tapping, which is expressed as: G G 1 / n i i U U n = = , (3) eq eq 1 , n i i C C = = (4) where UGi and Ceqi denote the induced voltage and equivalent capacitance of the ith cross section, respectively, and they are respectively calculated by (1) and (2). The horizontal and vertical geometric parameters of line configuration are required to calculate the equivalent circuit parameters. These geometric parameters can be easily found for the tower by referring to the design drawings, but they need to be solved for the mid span. The top view of a span is shown in Fig. 4. Since the phase conductors and shield wires are continuous at the mid span, the horizontal distances can be calculated by a trapezoidal approximation method. The horizontal distance between two wires at the mid span can be calculated by: 2 1 1( ) / ,y d d x l d= \u2212 + (5) where l represents the span (the horizontal distance between two hitch points), and d1 and d2 represent the exact horizontal distances in meters between the two wires at tower 1# and tower 2#, respectively. The vertical geometric parameters at the mid span denote the heights of the shield wires and the phase conductors above the ground" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001105_iros.2014.6943258-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001105_iros.2014.6943258-Figure2-1.png", "caption": "Fig. 2. Schematic of the model. It consisted of a rigid torso with a mass Mt and a foot mass Mf. Each leg contained a powered rotating hip with position \u03c6 and commanded torque \u03c4 as well as a powered prismatic leg with length l and commanded force F. The model was constrained to the sagittal plane.", "texts": [ " The exhaustive search results also show that, within the range of limb impedance values that result in stable locomotion, limb impedance does not significantly influence locomotion efficiency. The paper proceeds as follows: Section II describes the rigid body model, the locomotion controller, and the exhaustive search method. Section III presents exhaustive search results and joint work performed by individual simulation and robot trials. Section IV summarizes the findings and implications of the study. This study was performed on a simplified model of the MIT Cheetah robot. The joint layout and mass distribution of the quadrupedal model are shown in Fig 2. The model consisted of a rigid torso constrained to the sagittal plane with four legs of non-zero mass. Given that the action of the knee is to extend the foot relative to the hip, each limb was modeled as a rotating hip joint with a prismatic leg. This approach served to simplify the equations of motion without compromising the relevance of the simulation. The hip and knee joints were assumed to be ideal force sources with a maximum force and speed limit. The model parameters, shown in Table I, approximated the MIT Cheetah robot" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure4.13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure4.13-1.png", "caption": "Figure 4.13 Cylinder", "texts": [ " In practice, the circumference must take account of the material thickness. Any metal which is bent will stretch on the outside of the bend and be compressed on the inside. Unless the metal is of very light gauge, an allowance must be made for this. The allowance is calculated on the assumption that, since the outside of the bend stretches and the inside is compressed, the length at a distance half way between the inside and outside diameters, i.e. the mean diameter, will remain unchanged. The cylinder shown in Fig. 4.13 has an outside diameter of 150 mm and is made from 19 SWG (1 mm thick) sheet. Since the outside diameter is mean diameter = 150 \u2212 1 = 149 mm giving mean circumference = \u03c0 \u00d7 149 = 468 mm The circumference at the outside of the cylinder is \u03c0 \u00d7 150 = 471 mm Thus a blank cut to a length of 468 mm will stretch to a length of 471 mm at the outside and give a component of true 150 mm diameter. O b1 1 2 3 4 5 6 7 8 9 10 11 12 2 3 4 5 6 789 10 11 12a Circumference \u03c0d Figure 4.11 Development of cone O c ba Circumference \u03c0d Figure 4" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002408_s11012-020-01162-w-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002408_s11012-020-01162-w-Figure2-1.png", "caption": "Fig. 2 The flapping bird in three-dimensional overall motion", "texts": [ "2 Main body, the power generation components and the control systems The main body along with the power generation components and the control systems of the flapping mechanism can be modeled as a rigid body that has the translational and rotational movements in a three- dimensional space. To construct the corresponding bond graph of each subsystem, a body fixed coordinate axis that originates at the center of mass of the body, and its axes coincide with the main axis of the body is applied. According to Fig. 2, the angular and linear kinematic parameters of the body can be decomposed along the proposed coordinate axis. Using Newton\u2019s second law, the rotational and transitional relations can be extracted in the following way [30]: F \u00bc op ot rel \u00fex p \u00f025\u00de s \u00bc opJ ot rel \u00fex pJ \u00f026\u00de In the above equations F = [Fx,Fy,Fz] T and s = [sx,sy,sz] T, respectively are the vectors of net force and torque acting on the body. Moreover, p = [px,py,pz] T and pJ= [pJx,pJy,pJz] T, respectively are the linear and angular momentum vectors, x = [xx,xy,xz] T is the angular velocity vector, and t represents the time" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003277_tmag.2021.3053176-Figure11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003277_tmag.2021.3053176-Figure11-1.png", "caption": "Fig. 11. (a) 3-D-printed PLA molds for making SM2C cores and (b) as-fabricated cores from the SM2C in this work.", "texts": [ " To demonstrate the mold-ability of the tri-modal SM2C, we used it in a molding process to fabricate several magnetic core structures\u2014-C-core, E-core, toroid core, bar core, Authorized licensed use limited to: University of Prince Edward Island. Downloaded on May 15,2021 at 08:41:36 UTC from IEEE Xplore. Restrictions apply. DING et al.: SM2C WITH TRI-MODAL SIZE DISTRIBUTION FOR POWER ELECTRONICS APPLICATIONS 2800306 and plate core\u2014that are commonly found in power electronics circuits. First, we made the polylactic acid (PLA) molds shown in Fig. 11(a) by fused-deposition-modeling (FDM) 3-D-printing. Then, the SM2C paste was poured into the molds. After curing by the profile shown in Fig. 5, demolding, and light polishing, the magnetic cores were obtained and are shown in Fig. 11(b). Although in this work we used the molding process for making various cores, the SM2C paste can also be stored in a syringe and used as a feedstock in an extrusion-based 3-D printer to additively manufacture the cores [9], [30] or, as another example, to dispense the magnetics in a power electronics circuit on a PCB. We produced a SM2C that offers desirable magnetic properties for ease of making and integrating magnetic components in power electronics circuits. By combining three different magnetic powders with size distributions so chosen to increase the packing density, the tri-modal composite yielded a relative permeability over 30" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000530_jzus.a1900163-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000530_jzus.a1900163-Figure1-1.png", "caption": "Fig. 1 Geometric structure of the textured mechanical seal", "texts": [ " The model considers the effects of the Jakobsson-Floberg-Olsson (JFO) cavitation boundary condition, surface roughness, elastic-plastic contact, thermo-elastic deformation, and temperatureviscosity relation on the lubrication behavior. The study is conducted under the actual operating conditions of mechanical seals for aviation piston pumps. A comparative study of the three models for predicting seal performance, the HD model, THD model, and TEHD model, is performed. The distributions of the right triangular dimples are optimized under given operating conditions. The results provide new guidance for application of the pumping effect and present an optimum distribution of right triangular dimples. Fig. 1 shows the geometric diagram of a textured mechanical seal that consists of two rings. As shown in Fig. 1, pi is the atmosphere pressure and po is the fluid sealing pressure, ri and ro are the inner and outer radii of the sealing surface, respectively. The stator is fixed to the wall, while the rotor is a floating ring and spring loaded. Fspring is the spring force, \u03c9 is the rotational speed of the rotor. The sealing surfaces of the two rings are rough. The rotor is made of tin bronze and the stator is made of bearing steel (Table 1). The sealing surface of the stator, which has the greater hardness, is textured", " Thus, the heat flux produced by viscous shear and friction is considered to be entirely transmitted by conduction to the two rings (Luan and Khonsari, 2009). It can be considered as a local heat source on the sealing surface: 1 0, s T k q n (16) Yang et al. / J Zhejiang Univ-Sci A (Appl Phys & Eng) 2019 20(11):864-881 868 2 2 v f 0 con , r q q q f p r h (17) where n is the normal vector of the boundary, q is the heat flux, qv and qf are the viscous and frictional heat flux respectively, and f0 is the dry friction coefficient of the sealing surfaces. Thermal and mechanical boundary conditions can be seen in Fig. 1b. The surfaces of the two rings surrounded by sealed fluid are convective heat transfer boundaries. The surfaces of the two rings exposed to the air are adiabatic boundaries. The bottom surface of the stator is a fixed boundary and also an adiabatic boundary. The bottom surface of the rotor is pushed by a spring, and an axial displacement constraint is applied to its inside diameter. The convection transfer is conducted in the surfaces: 2 c 0( ) 0, s T k h T T n (18) where T0 is the bulk temperature of the fluid" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002507_s40435-020-00640-z-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002507_s40435-020-00640-z-Figure1-1.png", "caption": "Fig. 1 Model of the mechanism and kinematic definitions [5]", "texts": [ " Some brief explanations of controllers used in this work are given in Sect. 5. The performance of controllers applied on the compliant mechanisms in different scenarios are discussed and compared in Sect. 6. In Sect. 7, themost important results are presented and discussed and the conclusion is drawn. This section provides a short description of the discrete nonlinear modeling approach used for accurate modeling of the mechanical mechanisms based on the equivalent rigid link system (ERLS) concept [22]. As illustrated in Fig. 1, mechanism links are subdivided into finite elements. Furthermore, the entire motion of each complaint mechanism is splitted into large rigid-body motion of the ERLS and the small elastic deformation of the link. Particularly, the following notations are presented: \u2013 ri and ui are the position vectors of the nodes in the ith element of the ERLS and their elastic displacement respectively; these vector summation results in nodes motion of the i-th element in global reference frame. \u2013 pi is the vector of the position of a point of the i-th element \u2013 q is the ERLS generalized coordinate vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002769_j.procir.2020.03.091-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002769_j.procir.2020.03.091-Figure7-1.png", "caption": "Fig. 7: Predicted deformation results at P = 120 W and u = 750 mm/s.", "texts": [ " With further addition of third layer, the stress at interface raises to about 658.61 MPa (Fig. 5c). Furthermore, with the addition of fifth layer, the maximum residual stresses rapidly increase to 894.35 MPa (Fig. 5d). From the investigation, it has been found that the developed of residual stress on the fabricated parts are gradually increased with the successively adding product layers. The predicted distortion in printed component at laser power 120 W and scanning speed 1000 mm/s, 750 mm/s has been presented in Fig. 6, Fig. 7, respectively. At laser power 120 W and scanning speed 1000 mm/s, the predicted top four corner deformation from the numerical investigation were 205.2 \u00b5m, 213.4 \u00b5m, 263.5 \u00b5m, and 233.6 \u00b5m, as presented in Fig. 45. By decreasing the scanning speed to 750 mm/s (at same P = 120 W), the predicted top four corner deformation were increased to 239.5 \u00b5m, 246.3 \u00b5m, 311.8 \u00b5m, and 276.1 \u00b5m, as illustrated in Fig. 7. Experimental investigations were carried out to validate the numerical findings. Fig. 8 presented the experimental 3D printed part to measure the deformation of top four corner. The distortion was measured at the bottom right of the support material. The predicted top four corner deformation from the model are 248.7 \u00b5m, 257.7 \u00b5m, 332.1 \u00b5m, and 287.5 \u00b5m, which is in good agreement with the experimental measurement with the highest 74.2 \u00b5m deviation, as shown in Table 1. The difference between the prediction and experiment can be attributed to model simplification, including deformation relaxation in the experiment" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003746_s10409-021-01089-9-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003746_s10409-021-01089-9-Figure4-1.png", "caption": "Fig. 4 Details of the parabolic leaf spring model", "texts": [ " In the research published by Wang et\u00a0al. [47], the thickness of the leaf was estimated linearly. As introduced in previous section, the top and the bottom surface of the element are formulated individually. The users can set the components of the position and gradient vectors of the nodes on the top and bottom according to the data acquired from the geometrical description. Therefore, the parabolically changed thickness of the leaf can be modeled efficiently by the proposed SBE48, as shown in Fig.\u00a04. Spring eye is an important part of the leaf spring. It is reported that quite a lot of failure of the leaf spring happened on the spring eye. However, in most of the published continuum-based leaf spring model, the spring eye was omitted. The spring eye can be seen as a rubber bushing covered by steel cylinder. This type of structure is difficult to model by using LOBE24 because the nodes locates at two ends of the centerline. As a result, the element can only be assembled at the longitudinal direction", " Thereby the assembly of different components in a multi-body system can be efficiently achieved. If some of the linear constraints are released, specific mechanical joints can be described [41]. It also has been proved that even the gradient deficient element can work well with the ANCF-RF [52, 53]. In the research proposed by Tian et\u00a0al. [54], ANCF-RF was used to describe the rigid component in the elastohydrodynamic lubricated spherical joint. In this investigation, the rigid axle in the spring eye is 1 3 represented by an ANCF-RF, as shown in Fig.\u00a04. The nodes on the inner surface of the rubber bushing are fully constrained to the axle node. At the same time, the shackle is represented by two ANCF-RN, one shares the same position with the axle node, while the other is connected with the chassis. Pin joints are used to assemble these components, which are achieved by using linear constraint equations. The rigid-flexible model of the suspension system is thereby established. More details of the implementation of ANCF-RN could be found in the literature [41, 47, 52]", " It can be observed that with the increasing of the element number, the displacement curve is getting closer. In other word, the convergence property of the proposed element can be proved. Figure\u00a011 gives the configurations of the 2 \u00d7 2 \u00d7 2 mesh at different moments. The change of the kinetic, potential, elastic and the total energy of the 2 \u00d7 2 \u00d7 2 mesh is presented in Fig.\u00a012. As an isolated system, the total energy remains zero all the time. Therefore, the dynamic performance of the proposed SBE48 can be demonstrated. A leaf spring model introduced in Fig.\u00a04 is performed. Five ANCF-RNs are included in this model, among which one represents the chassis, two represent the axles and two represent the rigid shackle. Linear constraints are imposed between the axle nodes and the nodes inside the rubber bushing. Revolute joints are achieved also by the linear constraints to connect the chassis with the axles and shackles. The spring leaf has a flat bottom and the profile of the top and bottom surface of two wings are described by two cubic polynomials individually" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure1.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0000668_icsc47195.2019.8950675-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000668_icsc47195.2019.8950675-Figure3-1.png", "caption": "Figure 3. The FSMC diagram of the DFIG.", "texts": [ " Reactive powercontrol The sliding surface of the reactive power takes the following expression: s-ref sS(Q) = Q - Q (24) The derivative of equation (24) is: s-ref sS(Q) = Q -Q\u027a \u027a \u027a (25) By replacing the reactive power (9): drss-ref s M S(Q) = (Q - V I ) L \u027a \u027a \u027a (26) Then: s dr r dr s r qrs-ref s r M S(Q) = Q -V \u00d7(V -R I -g\u03c9 L \u03c3I ) L L \u03c3 \u027a \u027a (27) Hence the expression of Vdr = Vdr-eq + Vdr-n Replacing this expression in equation (27): s dr eq dr n r drs ref s r s r qr M S(Q) Q V ((V V ) R I L L g L I ) \u2212 \u2212\u2212= \u2212 \u22c5 + \u2212 \u03c3 \u2212 \u03c9 \u03c3 \u027a \u027a (28) During the sliding regime the state is stable therefore: S Q! = 0, \u027aS Q! = 0 et V + ) = 0 Which we deduce: 978-1-7281-1938-0/19/$31.00 \u00a92019 IEEE 467 . s r dr eq r dr s r qrs ref s L L V Q R I g L I ) MV\u2212 \u2212 \u03c3 = \u2212 \u2212 \u03c9 \u03c3 (29) and: dr-n vdr-nV = k \u00d7sign(S(Q)) (30) In the FSMC regulator, equations (23) and (30) become: dr-n vdr-nV = k \u00d7fuzzy (S(Q)) (31) qr-n vqr-n V = k \u00d7fuzzy(S(P)) (32) Fig.3 Illustrates the FSMC controller. V. SIMULATION RESULTS In order to conclude on the performance of the use of a fuzzy sliding mode control, we will present the simulations of the two commands (VC) and (FSMC) carried out on a DFIG. The block diagram of the simulations is presented by the figure (3) .with the parameters that are shown in Table. II A. Test without parametric uncertainty Figures 4 and 5 present the active and reactive powers produced by the DFIG with the different control strategies (CV and FSMC)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003547_j.jmapro.2021.04.060-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003547_j.jmapro.2021.04.060-Figure4-1.png", "caption": "Fig. 4. (a) Schematic of the laser scanning process, (b) schematic illustration of spot offset, and (c) overview and sideview of mesoscopic melting pool size and HAZ.", "texts": [ " Thereafter, the shrinkage/ warping of a representative rudder structure is employed to verify the corresponding scanning strategies. Contour and inner parameters correlate with the surface and mechanical properties, respectively. The contour parameters involve spot offset, scanning speed, and laser power. Inner parameters include laser power, scanning speed, and hatching space. The schematic of the laser J. Li et al. Journal of Manufacturing Processes 67 (2021) 195\u2013211 scanning process and parameters is described in Fig. 4(a). The melting pool (Fig. 4(c)) is obtained through the thermalrecrystallization-mechanical model. The lowercase w and d are the effective width and depth of the melting pool, which are indicated by the gray region with the temperature above 190 \u25e6C. To ensure the complete fusion of powders, the optimization criteria are w > h and t < d < 2t (h, t are the hatching space and layer thickness respectively, shown in Fig. 4 (a)). For the parameters of inter hatchings, the favorable depth is desired to be large than 1.5t. The heat-affected zone (HAZ, \u0394m in Fig. 4(c)) appears to be a culprit of the powder-absorption phenomenon of the layer border, which results in the dimensional inaccuracy of the printed components. The HAZ is the region with a temperature of 177 \u25e6C\u2013190 \u25e6C, which indicates the semi-melting state of powders. To minimize latent powder-absorption, the HAZ size is desired to be minimum. Fig. 4 (b) demonstrates the diagrammatic sketch of laser spot offset \u03b3. The border of the laser spot in red is tangent to the actual contour to diminish the dimensional deviation induced by overheating. Thus, the spot offset is equal to the half-width of the melting pool. Regarding the melting pool temperature, the maximum Tmax should be within the stable sintering range (SSR). SSR is the suitable temperature range to achieve successful sintering. It starts from the offset of melting temperature Tmf to the onset of decomposition temperature Tds" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001023_robio.2014.7090430-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001023_robio.2014.7090430-Figure5-1.png", "caption": "Fig. 5. Geometric model of Stewart platform.", "texts": [ " This section describes the PHARAD mechanism. We adopted Stewart platform, which is one of the parallel link mechanisms [19]. The moving plate and base plate are connected by six linear actuators, as shown in Fig. 4. The posture of the moving plate can be measured and controlled with six DOFs by measuring and changing the length of each linear actuator. By attaching a patient\u2019s foot to the moving plate, this mechanism can conduct ROM measurement and rehabilitation adapted to complex ankle motions. Fig. 5 shows the geometric model of the Stewart platform. and are the coordinate systems of the base plate (the upper plate) and the moving plate (the lower plate), and and are the origins of and , respectively. In , the position of is . The joint position where an actuator is connected to the base plate is (i = 1,\u2026,6), and the joint position where an actuator is connected to the moving plate is . In , the joint position where an actuator is connected to the moving plate is (i = 1,\u2026,6). In this study, we define the motions of the moving plate, which is attached to a foot sole, as the foot motions", " Analytically calculating the posture of the foot plate from the length of each actuator (forward kinematics) has not been done in this type of parallel link mechanism. Conversely, analytical calculations of the length of each actuator from the foot plate posture (inverse kinematics) are available. In this study, we numerically solved the forward kinematics using the inverse kinematics and calculated the posture of the foot plate , where is the 3D position ( ), and is the rotation angle ( ). A. Inverse Kinematics As shown in Fig. 5, can be expressed as , (1) where is a rotation matrix that rotates . From (1), the length of a link is calculated as . (2) From the above, when is known, can be easily calculated. When the length of each actuator is defined as , the relation between and can be determined as , (3) where is the Jacobi matrix, which is expressed as , (4) where (i = 1,\u2026,6) is the direction vector of the actuator, defined as (5) When the inverse matrix of exists, (3) can be converted into (7). , (6) . (7) is a measured value" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002972_s12555-020-0209-z-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002972_s12555-020-0209-z-Figure1-1.png", "caption": "Fig. 1. Simplified schematic of a rotary pendulum.", "texts": [ " The experimental analysis of the proposed controller is presented in Section 6. The paper is concluded in Section 7. The proposed adaptive controller is experimentally tested on a Rotary-Pendulum (RP) system. The RP is an open-loop unstable, highly nonlinear, under-actuated, and a multivariable electro-mechanical system [29]. These properties make it an ideal platform to study and validate the efficacy of the proposed adaptive controller [30]. The setup consists of a pendulum rod attached to a rotating arm, as shown in Fig. 1. The arm rotates with the aid of a permanent magnet DC motor. The motor\u2019s yaw motion is denoted as the arm-angle, \u03b1 . The pendulum\u2019s rod rotates about its pivot by using the energy provided to it via the arm\u2019s rotation. The angular displacement of the rod is denoted as the pendulum-angle, \u03b8 . The state-space model of a linear system is generally given by (1). x\u0307(t) =Ax(t)+Bu(t), y(t) =Cx(t)+Du(t), (1) where, x is the state-vector, y is the output-vector, u is the control-input signal, A is the system matrix, B is the input matrix, C is the output matrix, and D is the feed-forward matrix" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001203_s12541-014-0415-9-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001203_s12541-014-0415-9-Figure1-1.png", "caption": "Fig. 1 FZG back-to-back test rig", "texts": [ " Analytical method, namely the traditional gear strength analysis, is built on the basis of elastic mechanics, the gear contact strength calculation are based on the original Hertz formula through deformation and the coefficient correction. The results are more conservative and not accurate enough. Therefore, numerical method and test method were adopted for gear contact analysis in this paper. DOI: 10.1007/s12541-014-0415-9 The FZG rig is a back-to-back gear test rig of the closed power loop type, which is shown in Fig. 1. The actual test gears can be operated under very high loads. The driving motor only compensates for the power losses in the system, while the test gears can be operated under high load. The test rig consists of the drive gears and test gears, which are connected by two parallel shafts. The front shaft is split in two parts and carries the load clutch, necessary for the application of the load. The two flanges of the load clutch are twisted relative to each other and then bolted together for the load application" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001914_1.b35750-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001914_1.b35750-Figure8-1.png", "caption": "Fig. 8 Flow domain and boundary condition.", "texts": [ " Time-consuming computations should also be taken into account. Therefore, the sensitivity verification for the undetermined meshes was processed, and the results are depicted in Fig. 7. The variations of mesh quantity in each direction are proportional, so the total mesh quantity reflects the mesh quality in each direction. After comparison, we obtain a mesh with 2,469,000 elements for the final computational mesh of the straight-through seal and a mesh with 3,516,000 elements for the interlaced seal. Figure 8 shows the flow domain and boundary conditions of the CFD simulation. In Fig. 8b, the axial section is cut out from the flow D ow nl oa de d by W E ST E R N M IC H IG A N U N IV E R SI T Y o n A ug us t 5 , 2 01 6 | h ttp :// ar c. ai aa .o rg | D O I: 1 0. 25 14 /1 .B 35 75 0 domain. The boundary conditions of the inlet, outlet, and rotating wall are marked in the chart, and the uncharted boundaries are all no*slip walls. On account of the complete cycle model used in the computation, the periodicity boundary is not adopted in this research. After the rotor dynamics computation, the natural frequency of the rotor is solved as 247" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure20.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure20.9-1.png", "caption": "Figure 20.9 Portable crane", "texts": [ " the work is planned, organised and performed by competent people; Where appropriate, before lifting equipment and accessories is used for the first time, it is thoroughly examined. Lifting equipment may need to be thoroughly examined in use at periods specified in the Regulations (i.e. at least 6 monthly for accessories and, at a minimum, annually for all other equipments) or at intervals laid down in an examination scheme drawn up by a competent person. Following a thorough examination of any lifting equipment, a report is submitted to the employer to take any appropriate action. Small hydraulically operated portable cranes (Fig.\u00a020.9) are available which have an adjustable jib with a lifting capacity typically from 350 to\u00a0550 kg on the smaller models and up to 1700\u20132500 kg on the larger models. The smaller figure is the load capable of being lifted with the jib in its most extended position. In this case the D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 20 Moving loads 20 299 each hook shall be subjected to a proof load. After proof testing each hook is stamped to allow identification with the manufacturer\u2019s certificate of test and examination" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure17.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure17.12-1.png", "caption": "Figure 17.12 Location of punch in punch plate", "texts": [ " D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 17 Presswork 17 268 Another factor which must be considered in bending operations is the amount of springback. Metal that has been bent retains some of its original elasticity and there is some elastic recovery after the punch has been removed. This is known as springback. In most cases this is overcome by overbending, i.e. bending the metal to a greater extent so that it will spring back to the required angle \u2013 see also section 17.3.3. The punch is held in a punch plate and in its simplest form has a step as shown in Fig.\u00a017.12. The step prevents the punch from pulling out of the punch plate during operation. Length AB = 50 mm \u2212 inside radius \u2212 material thickness = 50 mm \u2212 5 mm \u2212 2 mm = 43 mm Length CD = 40 mm \u2212 5 mm \u2212 2 mm = 33 mm Since the radius is greater than twice the material thickness t, we can assume that the distance from the inside face to the neutral axis is 0.5t. \u2234 radius to neutral axis = 5 mm + (0.5 \u00d7 2 mm) = 6 mm Since it is a 90\u00b0 bend, length BC equals a quarter of the circumference of a circle of radius 6 mm \u2234 length BC = 2\u03c0R ____ 4 = \u03c0 R ___ 2 = 6\u03c0 ___ 2 = 9" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003279_j.matpr.2020.12.094-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003279_j.matpr.2020.12.094-Figure6-1.png", "caption": "Fig. 6. Stress distribution at three distinctive laser power (a) 25W (b) 50 W (c) 75 W.", "texts": [], "surrounding_texts": [ "The heat flux values for 25 W, 50 W and 75W were obtained from the ANSYS simulation for thermal analysis. The contour plots showing the variations in heat flux levels from minimum to maximum rise during ES-LBTM process are illustrated in Fig. 3. As ES-LBTM process occurred under submerged electrolyte environment; therefore, the transfer of thermal energy takes places among the source-medium-work material by means of heat transfer phenomenon which occurs from one atom to another. As at the higher laser power, the availability of heat energy was more thus, the impact for which the energy concentration levels escalated; which leads in rise of heat flux levels. Raise in heat flux levels signify more energetic atoms within the material, which depends upon the conductivity of the material and influenced by the heat energy at the point of laser-material interaction. From the thermal analysis for heat flux levels it has been observed that the maximum of 33.159W/mm2 at 25 W, 66.313W/mm2 at 50 W and 99.471W/ mm2 at 75 W are obtained. When the thermal energy is absorbed by the medium from the source to the work material, it results in the excess energy of the electrons and those excess energy is converted into kinetic energy of the free electrons and those electrons after gaining the excess energy tries to collide with the boundaries of neighboring atoms to transfer some of their energy to the lattice of the target material. These all are depended upon the material\u2019s response against the thermal energy, material\u2019s conductivity and the target spot area of laser interaction. As during the process the spot area is fixed corresponding to the beam diameter i.e. 0.021 mm, which leads to increase in the heat flux and the temperature at the target substrate. Fig. 4 and Fig. 5 illustrate the temperature disparities and mean temperature profiles (i.e. average temperature for material ablation) at the zone of laser beam and work body interaction. As we know that the DSS superalloy has melting temperature range of 1385 C \u2013 1443 C, but in simulation result it initiates to melt at 1060.3 C 1177.6 C with an average of 1146.4 C for 25 W, at 2093.5 C 2327.9 C with an average of 2265.5 C for 50 W and at 3124.4 C 3476 C with an average of 3382.3 C for 75 W which were displayed in dark yellow colour. Moreover, the material outsets to dissolute in the electrolyte at 1294.9 C 1412.2 C with an average of 1371.7 C for 25 W (refer to Fig. 5), at 2562.3 C 2796.7 C with an average of 2715.8 C for 50 W and at 3827.5 C \u2013 4179.1 C with an average of 4057.7 C for 75 Wwhich are shown in red colour. These temperature disparities relative to their average temperature for material ablation can be correlated from the Figs. 4 and 5 respectively. laser power (a) 25 W (b) 50 W (c) 75 W. From this endeavour it has been found that, with an increase in the laser power the genesis of temperature fields induces greatly; thus for which, it leads to achieve melting point at the target spot of material within less time interval. This also influenced the material ablation process to obtain the rate of ablation briskly with the specified hole feature due to the great disparities in temperatures. As ES-LBTM is a thermal process which surrounds with high thermal gradients (high heating and cooling rates), which leads to residual stresses associating with deformation. Moreover, eminently localized heating and cooling leads to uneven thermal inflation and shrinkage, which further reports in convoluted dissemination of residual stresses at the zone of interaction which leads to deformity across the neighbouring zone. From Fig. 1, we have observed that the laser current intensity impacts extremely as for laser power; the material ablation rate improved briskly. This is due to the reason for the raise in thermal energy, as thermal energy is directly proportional to the beam power. The raise in thermal energy further leads to induce more thermal shock waves at the spot of laser-material interaction, while laser passing through the electrolyte medium. Moreover, when the high intense beam of laser was passed through the electrolyte medium, the aggressive ions which were available within the solution lead to promote more thermally conductive atoms. The more number of conductive atoms with higher thermal energies bolsters to generate more thermal stress, which boosted the rate of material ablation. The stress, strain and total deformation obtained during ESLBTM approach are illustrated in Figs. 6, 7 and 8 respectively. From the analysis it has been observed that the effect of thermal stress is more at 75 W (i.e. 1187.9 MPa 10691 MPa) which leads to deform from 0.016198 mm to 0.14578 mm with an equivalent strain range of 9.1244e-15 to 0.053501, as shown in red colour in respective figures. However the induced thermal stress is less at 25 W (i.e. 397.68 MPa \u2013 3579.1 MPa) which leads to deform from 0.0054213 mm to 0.048792 mm with an equivalent strain range of 3.0728e-15 to 0.01791. From the Figs. 6, 7 and 8 it can be understood the material abalation rate at the zone of laser irradiation. It has been seen the great impact of stress and more values of stress in ES-LBTM approach. The presence of oxides in NaNO3 promotes the thermal process with some oxidation reactions at the zone of laser-electrolytematerial interactions. The temperature field distributions at 25 W, 50W and 75W from the ANSYS analysis is plotted with global minimum and maximum range in Fig. 9. This temperature distribution can be correlated with reference to the Fig. 4." ] }, { "image_filename": "designv11_22_0001065_j.asoc.2015.04.021-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001065_j.asoc.2015.04.021-Figure3-1.png", "caption": "Fig. 3. A simple example. A pendulum with length of 0.5 m has a mass M = 12 kg fi v", "texts": [ " 2, the proposed approach comprises three important fundamentals: dynamic inversion, robust performance metrics, and optimization with noisy parameters. In this work, the HPPOA algorithm is used for optimization because it has been considered competitive in the task of finding optimal or near-optimal global solutions for non-convex functions with noisy parameters [21]. Before describing how to design a tracking controller for unmanned aerial vehicles, Section 2.1 exemplifies the use of the proposed approach by using a simple example. 2.1. A simple example Fig. 3 displays a simple example of a nonlinear control system described in (1), where a pendulum turns around an horizontal axis when a torque T(t) is applied. The output variable is the angle (t), and the input variable is the torque T(t). \u0308 = 1 M \u00b7 L2 \u00b7 (T \u2212 M \u00b7 g \u00b7 L \u00b7 cos( ) \u2212 b \u00b7 \u0307) (1) Dynamic inversion consists of designing control signals that vary at the same rate as their respective state derivatives. This objective is accomplished by applying (2) to (1), where , \u0307 and c c \u0308c are used as a synthetic inputs", " The stimation of P based on N samples is (d) = 1 N \u00b7 N\u2211 k=1 F[S, H(qk), C(d)] (6) According to the optimization with noisy parameters, the PPOA algorithm is used as an optimization algorithm capable of inimizing (6), which is a functional with noisy parameters. In ummary, the process of optimizing (6) determines the gains of he controller that make the error go to zero. Simulation results are shown in Fig. 4. Simulations were caried out by considering \u00b130% of parametric uncertainty for all lant parameters shown in Fig. 3. The HPPOA algorithm carried ut 45 generations by using a population of 60 chromosomes and 0 Monte Carlo evaluations. nc = \u221a V2 \u00b7 ( \u0307c + PD )2 \u00b7 cos ( ) g2 Computing 34 (2015) 26\u201338 29 In this subsection, for simplicity reasons the theoretical second order system S was not optimized and the design was not formulated in terms of a multi-objective optimization problem. Now, a more detailed description about the proposed approach is presented in Sections 2.2\u20132.4. 2.2. Dynamic inversion The process of designing the controller begins by determining dynamic inversion and proportional-derivative (PD) layers, since DI is usually not applied alone" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000459_978-3-030-29041-2_16-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000459_978-3-030-29041-2_16-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_22_0000629_iecon.2019.8927706-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000629_iecon.2019.8927706-Figure2-1.png", "caption": "Fig. 2. One of the earlier LPTN [11].", "texts": [ " However, in this method, thermal resistances are calculated by using semi-empirical correlations preparing by some simplification assumptions or experiments may lead to uncertainty results [10]. This work was supported by the Estonian Research Council grant PUT (PUT1260). 978-1-7281-4878-6/19/$31.00 \u00a92019 IEEE 4372 One of the first studies in the field of thermal design of a totally enclosed fan cooled (TEFC) electrical machine is presented in[11]; where in 1991, Mellor developed the LPTN of the induction machine. To increase the accuracy of the model, he developed the LPTN in both axial and radial directions (Fig. 2) and finally validated it by doing the thermal test on two different induction machines in different power ranges from the medium (75-kW) to small power (5.5-kW). By calculating several convection heat transfer paths and developing the heat transfer paths in different directions, this model provides a very complex thermal network, which increases the computing process. In 2003, Boglietti [12] unveiled from his simplified thermal model for a TEFC induction machine and validated it by testing a 4-kW, 4-pole induction motor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000617_b978-0-12-814866-2.00007-5-Figure7.17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000617_b978-0-12-814866-2.00007-5-Figure7.17-1.png", "caption": "FIG. 7.17", "texts": [ " This is done by collecting a background spectrum without the sample and storing the result in the instrument\u2019s computer memory. The background spectrum is removed from the sample\u2019s spectrum by ratioing the two signals. In comparison to other instrument designs, an FTeIR provides for rapid data acquisition, allowing an enhancement in signal-to-noise ratio through signalaveraging. Therefore, Fast Fourier- Infrared (FT-IR) spectroscopy has been one of the powerful tool in material science research, in particular for identifying the prevailing functional group(s) and also monitoring reaction intermediates (Fig. 7.17). 7.11.4 Attenuated total reflectance (ATR) FTeIR instrument The analysis of an aqueous sample is complicated by the solubility of the NaCl cell window in water. One approach to obtaining infrared spectra on aqueous solutions is to use attenuated total reflectance instead of transmission. Fig. 7.18 shows a diagram of a typical attenuated total reflectance (ATR) FTeIR instrument. The ATR cell consists of a high refractive index material, such as ZnSe or diamond, sandwiched between a low refractive index substrate and a lower refractive index sample" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000709_psc49016.2019.9081455-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000709_psc49016.2019.9081455-Figure4-1.png", "caption": "Fig. 4. set of windings attached to the sides of train [3]", "texts": [ " Electrdynamic suspension Unlike electromagnetic suspension that is based on absorption force, Electrodynamic Suspension (EDS) is based on the electromagnetic repulsion force produced between a variable field produced by the train and a set of coils which are attached on the track. Fig. 3 demonstrates the basic layout of this type of suspension [1]. In practice, a set of windings (usually Superconductive coils), in which DC current flows, are attached to the sides of train while an other set of coils in shape of 8, which are short circuited (Fig. 4) are attached on the sides of the track. It can be shown that, the currents produced in the track coils, tends to make the magnetic flux seen by the upper and lower loops equal and keep the train on the center of the 8-shaped coils [3]. Compared to EMS, this system has got a better stability in higher speeds, but two major problems are associated with it. First of all, in order to be in the superconductive state, superconductors must be cooled down to about 5\u00b0K, which requires a complex cooling system and secondly, in lower speeds the suspension force is reduced and therefore the train needs to rely on wheels for speeds lower than 100 km/h" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001221_esda2014-20001-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001221_esda2014-20001-Figure6-1.png", "caption": "FIGURE 6. TYPICAL DESIGN OF THE COAXIAL NOZZLE. (SOURCE [4])", "texts": [], "surrounding_texts": [ "Conventional industrial laser cladding equipment is composed of the laser source, requisite optical components, processing head, powder or wire feeder as well as guidance and processing software. In the case of the in-situ repair of marine crankshaft journals, the laser nozzle and powder supply equipment has to fulfil particular requirements: very good accessibility; suitability for sustained processing, including resistance under high thermal impacts (e. g. caused by back-reflection of the laser beam onto the work piece surface); highly efficient powder supply; flexible and rapid integration of the laser cladding equipment; capability for omni-directional movements during processing. Since the aforementioned technology is in its primary development stage, emphasis in this article is given to the selection of the cladding nozzle. Other, equally important elements of the cladding equipment are ignored at this point and will be addressed in subsequent stages of the research. The type of nozzle, the angle of the powder stream, the powder profile in the process zone and powder stream diameter in the melt pool area all influence the interaction of powder particles with the surface. An appropriate nozzle is one that provides the minimum solid particles with solid surfaces. Minimising the impact between the solid particles and solid surfaces increases the powder catchment efficiency [4]. Current industrial cladding powder nozzles are designed predominantly with coaxial powder delivery systems (see Fig. 7). These nozzles are designed to perform build-up process in all directions. Use of nozzles with powder supply today accounts for 60 to 80% of actual industrial applications. Research would indicate that up to 98% of conventional tasks can be performed with coaxial nozzles, although they always require perfect conditions, i.e. a melt pool diameter of more than 5 mm [6, 13]. The use of an internal mixing chamber means that up to four different powdery materials can be simultaneously Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 7 Copyright \u00a9 2014 by ASME delivered and mixed. Typically a wide range of metals (e.g. Fe-, Ni- or Co-based), but even carbides (e.g. WC, VC, TiC) as well as ceramics can be used. The described nozzle systems are suitable for production needs with up to 10 kW of laser power. Typically these nozzle systems are designed with an integrated ring slit for smooth, stable powder delivery. In some cases however, a high tilting angle of the cladding equipment in the crankshaft housings might even require the use of discrete nozzle systems. Instead of a ring slit at the nozzle tip, separate powder channels enable greater tilting of the processing head. Furthermore the available space of the coating area might even yield modifications of the nozzle design. By a rearrangement of the internal media supply and cooling channels as well as laser beam arrangements, the nozzle system can be made very slim in order to achieve greater accessibility (see Fig. 8). In this case, a partially reduced deposition rate should be considered. Despite the high efficiency of powder, not all powdery materials can be placed on the crankshaft. Hence, the wirebased weld metal delivery with full filler material utilization can offer an alternative to the powder process. Certain wire cladding heads even enable repairs in omni-directional mode." ] }, { "image_filename": "designv11_22_0003510_s12206-021-0424-4-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003510_s12206-021-0424-4-Figure14-1.png", "caption": "Fig. 14. Miura-ori membrane deployment simulation.", "texts": [ " In the deploying process of membrane, the magnitude of the force on each corner in the x direction ranges from 0 to 10.341 N. Beyond this range, the model calculation does not converge. When the force in the x direction of each corner exceeds 10 N, the membrane tends to be completely flat. In addition, when the load continues to increase, the deployment rate and the deployment error change little. So take 10 N as the peak value of each corner force load in the x direction, and the specific values in each direction are shown in Table 2. Fig. 14 shows the deploying process of membrane. As shown in Fig. 14(a), because the force of the area of the crease intersection is more complicated, when the model is in the initial configuration there is still some elastic deformation in some areas of the membrane. The von Mises stress is the largest at the intersection of the creases, and it is 24.27 MPa, and gradually decreases to 0.2 MPa as the creases extend outward. Different from the force drive, the speed load is defined as a constant value during the deploying process of the membrane. According to the displacement in the x and y directions between the initial configuration and the deployed configuration of membrane corners by force driving, the specific values of the constant-speed are calculated, as shown in Table 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003053_iros45743.2020.9341428-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003053_iros45743.2020.9341428-Figure6-1.png", "caption": "Fig. 6: Example of robot team cooperation as valid assembly solution found by the planner.", "texts": [ " Although the high complexity of the planning problem, the overall planning time is remarkably good and still just a fraction compared to the execution time. For building the bridge, a suitable set of robot cooperation is required to guarantee a non-collapsing assembly. A key moment of assembling the bridge is the placement of the uppermost brick of the arch which first connects both side parts. Here, the planner has to find a coordinated application of the right skills that never leave parts of the bridge unsupported (cf. Fig. 6). In the situation described by (a), two robots R1 and R2 are supporting each one part of the bridge at the second topmost brick. Robot R3 is not able to place the final connection brick as it would cause the directly underlying bricks to break. Thus, robot R3 moves in position to support one of these upper bricks (b), which enables R1 to release its support on this side (c). Repeating this scenario with Robot R1 on the other topmost brick (d), R2 is free to release its support (e) and to care about the critical connection brick which can now safely be placed (f)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001240_icuas.2014.6842360-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001240_icuas.2014.6842360-Figure1-1.png", "caption": "Fig. 1. Pure pitching motion", "texts": [ " Finally, section VI presents the conclusions and future work. In order to obtain the model equations, by omitting any flexible structure of the MAV, the fixed-wing MAV is then considered as a rigid body. Also we do not consider the curvature of the earth, it is considered as a plane, because we assume that the fixed-wing MAV will only fly short distances. With the previous considerations, we obtain the model by applying the Newton\u2019s laws of motion. The parameters involved in the longitudinal dynamic model (1)-(5) are shown in Figure 1. These parameters allow to analyzing the movement toward the front of an airplane [5], particularly the altitude control. V\u0307 = 1 m (\u2212D + T cos\u03b1\u2212mg sin \u03b3) (1) \u03b3\u0307 = 1 mV (L+ T sin\u03b1\u2212mg) sin \u03b3) (2) \u03b8\u0307 = q (3) q\u0307 = M Iyy (4) h\u0307 = V sin(\u03b8) (5) where V is the magnitude of the airplane speed, \u03b1 describes the angle of attack, \u03b3 represents the flight-path angle and \u03b8 denotes the pitch angle. In addition, q is the pitch angular rate (with respect to the y-axis of the aircraft body), T denotes the force of engine thrust, h is the airplane altitude [5] and \u03b4e represents the elevator deviation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001221_esda2014-20001-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001221_esda2014-20001-Figure3-1.png", "caption": "FIGURE 3. IN-SITU LASER CLADDING DEVICE\u2019S GUIDEWAYS AND FRAME", "texts": [ " The damaged crankshaft journal surface is renovated by fitting the laser cladding nozzle positioning and guidance device directly on the crankshaft journal fillets. These fillets as a rule are not damaged or worn and thus conserve the original manufacturers\u2019 crankshaft dimensions (see Fig.2). The internal fillets can therefore be used as a reference base, enabling the laser cladding nozzle guidance platform to be positioned using the device\u2019s specially designed guide-ways. Drawings of the \u0160koda/\u010cKD type \u201c6-27.5 A2L\u201d medium-sized marine diesel engine crankshaft fillets in Fig. 3 are provided as an illustrative example. In this case, the diameter of the crankshaft main bearing is 210 mm with internal fillet (see Y in Fig.3) radius of 11 mm, and the crankpin journal is 190 mm with internal fillet radius of 12 mm (see X in Fig.2). Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 5 Copyright \u00a9 2014 by ASME The essence of the research is a device for the repair and renovation of crankshaft journal surfaces. Its technical description is provided below to highlight the complexity and challenges as well as the clear benefits of this endeavour. The device comprises two guide-ways and two opposite guide-ways for positioning the device on crankshafts fillets and two frame parts, each of which is fixed to the respective guideway. The device also comprises two upper rods, positioned in the upper part of the frame part, and two lower rods in the lower part of the frame part, by means of which both frame parts are rigidly connected to each other (see Fig. 3). The device further comprises two carriages which are installed on the upper rods and lower rods so that both carriages can be slid along these rods. A laser nozzle is installed operatively between both carriages. The device includes two servo motors, the first of which is installed in the first carriage and operatively connected to the laser nozzle to control its pivoting angle (see Fig. 4). A second servo motor is installed on the second carriage and is operatively connected to one of the two lower rods by means of a gearing transmission to control the laser nozzle\u2019s longitudinal position", " Two supporting plates are permanently fixed on the opposite guideways. Furthermore, these supporting plates are connected to each other by two opposite-rods by means of which both perpendicular guide-ways are in fixed connection to each other. When installed on the crankshaft journal, the guide-ways and opposite guide-ways are connected and fixed to each other by means of four adjustable arms. The adjustable arms are connected to the guide-ways and opposite-guide-ways by eight guidance-screws (see Fig. 3). While the crankshaft is being rotated around its main axis, the laser head top-down position is ensured by eight pneumatic cylinders. These cylinders are connected to the guide-ways and opposite-guide-ways by the aforementioned eight guidancescrews. The pneumatic cylinders can freely rotate around these eight guidance-screws. Downloaded From: http://proceedings.asmedigitalcollection.asme.org/ on 01/30/2016 Terms of Use: http://www.asme.org/about-asme/terms-of-use 6 Copyright \u00a9 2014 by ASME The crankshaft journal surface repair technology using the aforementioned device involves the following steps: a) positioning the device above a damaged (renewable) journal surface by means of the guide-ways and opposite guideways; b) applying a cladding powder onto the damaged journal surface; c) positioning a laser nozzle above the damaged journal surface by means of two servo motors; d) irradiation of the cladding powder by a laser beam emitted from the laser nozzle; e) repeatedly performing steps b) through d) until the damaged journal surface is cladded" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure13-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure13-1.png", "caption": "Fig. 13. Stress analysis results of the internal gear with single contact point (unit: MPa): (a) von Mises stress; (b) shear stress.", "texts": [ " The maximum stress occurs at contact point, which is located on the middle of the tooth profile. It has regular elliptical distribution along the direction of tooth width, and the distribution area has the trend of expanding to the tooth root direction. With the increase of the contact area, the contact stress will gradually decrease. The maximum von Mises stress and shear stress of the pinion with single contact point in Fig. 12 are 729.02 MPa and 374.89 MPa, respectively. The maximum von Mises stress and shear stress of the internal gear with single contact point in Fig. 13 are 729.58 MPa and 201.07 MPa, respectively. 4.2 Stress analysis with two contact points Similarly, we established a gear model with two contact points according to the proposed generation methods of tooth profiles. It means that there are two contact points on one contact pair at the same time during the meshing process. Here, the meshing pair should also be the pinion with convex tooth profile and the internal gear with concave tooth profile. According to the derived conclusions in Sec. 2, we can obtain tooth surfaces of internal gear pair with two contact points as follows" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001813_978-94-017-7515-1_19-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001813_978-94-017-7515-1_19-Figure1-1.png", "caption": "Fig. 1 The micro-structure of a flax fibre cell (Baley 2002), reproduced with permission from Elsevier", "texts": [ " Flax fibres are produced in the stems of flax bast plant and are a cellulose polymer like cotton fibres. However, its structure is more crystalline which makes it stronger than cotton and it is stiffer to handle and is more easily wrinkled. Length of a flax plant extends to 100 cm which has strong fibres along its stem with average fibre diameter 10\u2013 25 \u00b5m (Bos et al. 2002). The micro-structure of a flax fibre is complex due to the hierarchical organisation at different length scale and different materials present in variable proportions (Fig. 1). From the Fig. 1, it can be seen that the thickest cell wall is S2. This wall contains numerous crystalline cellulose micro-fibrils and amorphous hemicellulose which are oriented at 10\u00b0 with the fibre axis that gives the high tensile strength to the flax (Baley 2002). Flax fibre consists of cellulose, hemicellulose, wax, lignin and pectin in different quantities were reported by many authors (Bastra 1998; Troger et al. 1998; Lilholt et al. 1999). This variation of proportions in the constituents of flax fibres is due to the plant variety, agriculture variables i", " Because of its eco friendly production, we hope that flax fibre can find a better place for itself in the composite arena as they are a good candidate for future generation of bio-fibres, especially in automotive industry; because natural fibres possess excellent sound absorbing efficiency, are more shatter resistant and have better energy management characteristics than glass fibre reinforced composites. Therefore in future research, flax composites\u2019 durability, moisture resistance, surface modification techniques and the mechanical properties can be improved. Acknowledgments The authors are grateful to Elsevier Publishers and authors who permitted to re-produced Fig. 1 and Table 1 from their publications. Also, the first author wishes to thank for the financial support for this research which was provided by the Scientific Research Project Unite of Marmara University project grant FEN-D-110315-0071 of 2015. Pavithran C, Mukherjee P.S, Brahmakumar M, Damodaran A, Impact properties of natural fibre composites, Journal of Materials Science Letters, 1987, 6, 8 p.882-884. Mohanty AK, Misra M, Drzal LT, Selke SE, Harte BR, Hinrichsen G. In: Mohanty AK, Misra M, Drzal LT, editors" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001875_978-4-431-54595-8-Figure2.7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001875_978-4-431-54595-8-Figure2.7-1.png", "caption": "Fig. 2.7 (a) Experiment setup, top view. Subjects were asked to produce hand force along eight directions in order, from direction 1 to direction 8, under an isometric condition. The task was performed with a left/right arm, and it was identical for both arms. (b) Sketch of examined eight muscles which mainly contribute to the studied movement", "texts": [ " The subjects of the experiment are three healthy volunteers (all of whom are 22 years old, male, and right handed) with no record of neuromuscular deficits. The experimental procedures were conducted with the approval of the ethics committee at Osaka University. The subjects were asked to produce 8N of hand force pointing along eight directions under an isometric condition in a horizontal plane, starting at direction 1 and shifting orientation by 45\u0131 in the counterclockwise rotation until ending at direction 8 (Fig. 2.7a). The hand position was 0.28 m in front of the subject\u2019s trunk. The task goal was to produce hand force as close to the reference force as possible and to keep maintaining the force in a 6-s duration before moving to the next direction. The task was performed with the left hand and then with the right hand, and it was identical for both hands. The subjects sat comfortably on chairs, with the elbow supported to reduce the gravitational effect and to allow for shoulder and elbow flexion-extension in a horizontal plane. The subjects grasped a joystick and pulled or pushed it to produce hand forces at a comfortable speed while looking at reference forces displayed on a screen. For simplicity, wrist movement was prevented by a splint so that it could be ignored in 38 H. Hirai et al. the analysis. The EMGs of eight muscles (see Fig. 2.7b) that mainly contribute to the studied task were collected by using a multi-telemeter system (WEB-5000, Nihon Kohden Corp., Japan) at 1000 Hz. Examined muscles were identified according to the guidelines in Hislop and Montgomery (2007). After cleansing the skin to reduce the resistance below 10 k , surface electrodes were placed on the examined muscles. EMG signals were band-pass filtered (0.03\u2013450Hz), hum filtered (60 Hz), amplified ( 2000), and stored in a computer. A force sensor (USL06-H5-200, Tech Gihan Co" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001563_detc2015-46601-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001563_detc2015-46601-Figure1-1.png", "caption": "Figure 1. Test apparatus used for bearing parameter identification", "texts": [ " In this paper, linear model and three kinds of nonlinear oil-film force models are derived and the linear as well as nonlinear dynamic coefficients are identified based on the experimental data. In this investigation, a method based on least squares estimation in frequency domain is presented. All experiments were executed on a suitable test rig with a five-pad TPJB of 100 mm diameter. TEST RIG AND BEARING DESCRIPTION A thorough description of the test rig used in this work is given in [26]. The load is applied in the middle of the shaft by means of two hydraulic actuators (pos. I in Figure 1) placed in an orthogonal configuration at \u00b145 degrees with respect to the load cells. The actuators are connected to the shaft by two deep groove precision ball bearings. The hydraulic actuators have a nominal force of 25kN and are able to displace the shaft with amplitude of 0.1mm with a band of 0-50Hz and are provide by high resolution position and force transducers. All data are sampled by PC-based control software with Labview software through PCI DAQ boards with model of NI cDAQ-9178. The photo of real test rig is shown in Figure 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000514_j.mechmachtheory.2019.103667-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000514_j.mechmachtheory.2019.103667-Figure1-1.png", "caption": "Fig. 1. Schematic representation of a face-gear drive with supporting shafts.", "texts": [ " \u2022 Investigation of the influence of shaft deflections on the formation of the bearing contacts, and maximum contact and bending stresses. \u2022 Development of a computational approach for the determination of the geometry of face-gear tooth surfaces by adjusting the relative position of the shaper cutter to compensate the relative errors of alignments caused by shaft deflections when a nominal torque is applied. \u2022 Development of an alternative and simplified method for the compensation of shaft deflections easy to implement in actual cutting machines. Fig. 1 shows a schematic representation of a face-gear drive with supporting shafts. Points P1 and P2 are located at the intersection between the pinion axis of rotation and its corresponding front and back planes, respectively, whereas points W1 and W2 represent counterparts of points P1 and P2 for the face gear. The shafts of pinion and face-gear are designed as hollowed with outer and inner diameters represented by do1 and di1 for the pinion and by do2 and di2 for the face-gear (see Fig. 1). Points B1, B2, B3 and B4 are located at the position of the supporting bearings for the shafts. Both gears are mounted in overhanging configuration. The transmitted nominal torque of 2500 N m is applied at B2 on the pinion shaft, and the face-gear shaft rotation is restricted at B4. Fig. 2 shows all the components of the finite element model of the face-gear drive for determination of errors of alignment due to shaft deflections. Here, five pairs of contacting teeth have been used to avoid influence of the boundary conditions on the results and to take into account the load sharing between pinion and face-gear tooth surfaces" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003453_s12206-021-0334-5-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003453_s12206-021-0334-5-Figure2-1.png", "caption": "Fig. 2. (a) Proposed new serial chain, distributed compliant mechanism; (b) length comparison of compliant five bar and compliant serial chain mechanism.", "texts": [ " Also, since the mechanism has 2 DOF, tip coordinates should be detected accordingly using two-dimensional position sensor however this would increase the cost of the experimental setup. Tip trajectory control of the proposed mechanism can be attained in an open loop manner such that if the reference angles are calculated for a well-defined trajectory using the inverse kinematics, then it\u2019s expected that the tip would follow the same trajectory. Relating the lateral and axial tip coordinates of the proposed fully compliant, closed chain mechanism as depicted in Fig. 2(a) to the rigid link angles yields two equations with three unknowns. Since the number of unknowns is more than the number of equations from the geometric constraints, problem can\u2019t be solved. On the other hand, for the initial compliant five bar design shown in Fig. 2(b), rigid links on each side of the mechanism can be calculated using the analytical kinematic analysis along with the initial angular velocities. We created two experimental setups as seen in Fig. 11 and applied the same inputs to the servos and observed that the tip follows the same trajectory if the horizontal distance between actuators were adjusted since the overall geometry of both mechanisms is same. Therefore, we adopted the simplified model of the compliant five bar mechanism [25] to control the tip of the closed chain, fully compliant mechanism for the ease of fast response in real time" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002092_0954406216682768-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002092_0954406216682768-Figure4-1.png", "caption": "Figure 4. Normal contact force and deformation for a ball.13", "texts": [ " The synthetic curvature of the contact point between the ball screw and the ball and the synthetic curvature between the nut and the ball is expressed as follows13,19 X 1 \u00bc 2 db \u00fe 2 db 1 dbf \u00fe 2 cos cos d0 db cos \u00f02:1\u00de X 2 \u00bc 2 db \u00fe 2 db 1 dbf 2 cos cos d0 \u00fe db cos \u00f02:2\u00de where d0 is the nominal diameter of the ball screw; db is the diameter of the ball; is the pressure angle between the ball and groove; f is the dimensionless radius of curvature of the groove; is the helix angle of the ball screw. Hertz contact theory represents the relation between normal force and elastic deformation along the force direction. Based on Figure 4, the axial contact deformation a, which is caused by the normal contact deformation of the rolling ball can be expressed as follows. a \u00bc 1 \u00fe 2 sin cos \u00f03\u00de The equation between axial thrust and axial contact deformation of a ball can be established as equation (5).13 a \u00bc Ke1 \u00fe Ke2\u00f0 \u00deF2=3 bx \u00f04\u00de where Fbx is the axial thrust of a ball. The contact coefficient Kei can be expressed as follows Kei \u00bc Ji mai 3 2 1 u21 E1 \u00fe 1 u22 E2 2( )1=3 1 sin cos \u00f0 \u00de 5=3 X 1=3 i \u00bc 1, 2 \u00f05\u00de The axial contact force of the nut is Fa \u00bc Pz sin cos \u00bc zFbx \u00f06\u00de where z is the number of the working balls in the nut" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002319_j.jsg.2020.104023-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002319_j.jsg.2020.104023-Figure10-1.png", "caption": "Fig. 10. Folds and boudins in bulk plane strain. The distribution of possible fold and boudinage axes is shown in relation to lines of no finite and no infinitesimal deformation in equal-area net (after Flinn, 1962). X > Y > Z \u00bc axes of principal strain; X-axis is vertical in the equal-area net. (a) Undeformed state showing the initial orientation of competent layers used for the different runs. Great circles of competent layers are shown by different colors. Numbers in black boxes indicate the initial angle between layer and Z-axis (\u03b8Z(i)). (b) Deformed state showing the position of passive planes that were initially oriented like the competent layers used in the present study. (c) Deformed state showing the position of competent layers of deformation series 1 (m \u00bc 82). (d) Deformed state showing the position of competent layers of deformation series 2 (m \u00bc 18). The color of the great circles shown in (b)\u2013(d) corresponds to the color used for the great circles shown in (a). Numbers in black boxes indicate the initial angle between passive plane (or competent layer) and Z-axis. For further explanation, see text.", "texts": [ " As the rate and amount of layer rotation are significant for the degree of structural asymmetry and for the strain the layer undergoes with progressive deformation, this topic has to be particularly considered when discussing the results presented above. The rate and amount of layer rotation depend on the bulk finite strain (eZ), the initial attitude of the layer (\u03b8Z(i)) and the viscosity ratio between layer and matrix (m). When the lines of no finite deformation and the lines of no infinitesimal deformation for a plane-strain ellipsoid are plotted on an equal-area net, they divide the net into three different areas (Flinn, 1962, Fig. 10a). All directions whose poles lie inside the line of no finite deformation (thin dotted line in Fig. 10a) have been elongated, while all directions lying outside have been shortened. The lines of no infinitesimal deformation (dashed line in Fig. 10a) divide the area outside the line of no finite deformation into two areas. Directions lying between the two lines have at first been shortened and later elongated, but insufficiently to restore their original lengths. Directions outside the line of infinitesimal deformation have all been shortened. The three areas, shown in red, pink, and blue in Fig. 10a, are called the area of elongation, the area of reduced shortening and the area of shortening (Flinn, 1962). In the course of a single coaxial deformation, a direction may first undergo shortening until the line of no infinitesimal deformation is crossed, then elongation until it has regained its initial length at the line of no finite deformation, and finally absolute elongation (Ramberg, 1959; Flinn, 1962). If one principal direction in a plane is a direction of shortening, then the other principal direction in that plane is a potential fold axis. On the other hand, if one principal direction is a direction of extension, then the other is a potential boudinage or pinch-and-swell axis. In the present case of coaxial plane strain, single boudinage axes will develop in the field of shortening and reduced shortening, whereas single fold axes will develop in the field of elongation. The fields of boudinage and fold axes are depicted in Fig. 10a, which also shows the initial attitudes of the competent layers used in the present study. To show the difference in rotation between the competent layer and a corresponding passive plane, a sphere (circle) with inscribed passive planes (lines) was incrementally deformed until the final shortening strain of the present experiments (eZ \u00bc 70%) was attained (Fig. 11). Initially, the attitude of planes (lines) of no finite strain and of no infinitesimal strain is the same at \u03b8Z \u00bc 45\ufffd. The planes (lines) of no infinitesimal strain will remain at this position during the entire deformation, whereas oblique passive planes rotate around the Y-axis", " Passive planes in the field of shortening, reduced shortening, and elongation are shown in blue, pink, and red, respectively. It is obvious from Fig. 11 that with increasing strain, the number of planes (lines), which are undergoing shortening or reduced shortening, decrease, whereas the number of planes undergoing elongation increase. The angles between the Z-axis and the rotated passive planes (\u03b8Z(RPL)) are listed in Table 1 and are added to the diagrams shown in Figs. 6a and 9a. The final position of the passive planes is depicted in Fig. 10b. The final position of the competent layers for m \u00bc 82 and m \u00bc 18 is shown in Fig. 10c and d, respectively. In cases where \u03b8Z(i) was <45\ufffd, rotation of the competent layer was particularly retarded compared to the rotation of a corresponding passive layer (Figs. 6a, 9a and 10a-d). As the layers are folded in these cases, the geometry of F1 and F2 folds might have affected the rate and degree of overall rotation of the folded layer. Moreover, the rotation angle is not always well constrained because of the folding. This problem is not given if \u03b8Z(i) was >45\ufffd. In these cases, the layer was free from significant folding, but underwent either boudinage or only homogenous thinning", " F1 and F2 folds result in Typ 0 fold interference patterns, which show the lowest degree of interference of superposed folds since F1 parasitic folds remain almost cylindrical and plane and there is no real pattern of interference in cross sections of such folds (Ramsay, 1967, p. 531; Thiessen and Means, 1980). However, migration of F1 hinges might have occurred during F2 folding, which could explain the striking difference in length of short and long limbs of F1 folds when approaching a F2 hinge (Fig. 8b). Similar superposed folds are present in multilayers of naturally deformed rocks (e.g. Schmalholz and Mancktelow, 2016: Fig. 10). These J. Zulauf et al. Journal of Structural Geology 135 (2020) 104023 J. Zulauf et al. Journal of Structural Geology 135 (2020) 104023 J. Zulauf et al. Journal of Structural Geology 135 (2020) 104023 parasitic folds, however, result from contact strain (Ramberg, 1961, 1963, 1964). The matrix between competent layers shows a deformation close to pure shear in the hinge area and a combination of pure and simple shear in the limb areas (Frehner and Schmalholz, 2006). Passive planes oriented at \u03b8Z(i) \u00bc 10\ufffd\u201315\ufffd started with shortening and were subsequently elongated due to rotation; however, they remained in the field of reduced shortening until the final strain of eZ \u00bc 70% was attained (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001015_chicc.2014.6896862-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001015_chicc.2014.6896862-Figure2-1.png", "caption": "Fig. 2: Communication topology: L stands for the leader (or reference signal) and F1, \u00b7 \u00b7 \u00b7 , F4 stand for 4 manipulators.", "texts": [ " The dynamics of this two-link manipulator can be described by the following EL equation [20] M(q)q\u0308 + C(q, q\u0307)q\u0307 +G(q) = \u03c4, (15) where q(t) = (q1(t), q2(t)) T (\u03b81(t), \u03b82(t)) T , M(q) = [ M11(q) M12(q) M21(q) M22(q) ] , C(q, q\u0307) = [\u2212hq\u03072 \u2212hq\u03071 \u2212 hq\u03072 hq\u03071 0 ] , M11(q) =m1l 2 c1 + I1 +m2(l 2 1 + l2c2 + 2l1lc2 cos(q2)) + I2, M12(q) =M21(q) = m2l1lc2 cos(q2) +m2l 2 c2 + I2, M22(q) =m2l 2 c2 + I2, h =m2l1lc2 sin(q2), G(q) =(g1(q), g2(q)) T , G1(q) =m1lc1g cos(q1) +m2g[lc2 cos(q1 + q2) + l1 cos(q1)] G2(q) =m2lc2g cos(q1 + q2); and g is the gravitational acceleration, \u03c4 \u2208 R 2 is the input torque, Ii is the moment of inertia of linki. The parameters setting are shown in the Table 1. The communication topology is illustrated in Fig. 2. The Laplacian matrix is LG = \u23a1 \u23a2\u23a2\u23a2\u23a2\u23a3 0 0 0 0 0 \u22121 3 \u22121 0 \u22121 0 \u22121 1 0 0 0 0 \u22121 1 0 0 \u22121 0 \u22121 2 \u23a4 \u23a5\u23a5\u23a5\u23a5\u23a6 . (16) Hence, L2 = \u23a1 \u23a2\u23a2\u23a3 3 \u22121 0 \u22121 \u22121 1 0 0 0 \u22121 1 0 \u22121 0 \u22121 2 \u23a4 \u23a5\u23a5\u23a6 , (17) whose eigenvalues are {0.2222, 3.9230, 1.4274\u00b1 0.5071i}. Therefore \u03c3min = 0.2222. The leader\u2019s trajectory is q0(t) = [\u221260 0 ] + [\u221211 \u221216 ] t+ [\u22120.6 0.8 ] t2+ [ 0.008 \u22120.01 ] t3. (18) The protocol is \u03c4i(t) = M(qi) 5\u2211 j=0 \u03b1ij ( k1(q\u03070(t)\u2212q\u0307i(t))+k2(qj(t)\u2212qi(t)) + k3 \u222b t 0 (qj(s)\u2212 qi(s))ds + k4 \u222b t 0 \u222b s1 0 (qj(s2)\u2212 qi(s2))ds2ds1 ) + C(qi, q\u0307i)q\u0307i(t) +G(qi), (19) where K = (k4, k3, k2, k1) = max{1, \u03c3\u22121 max}BTP = (4" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001113_1.4029624-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001113_1.4029624-Figure2-1.png", "caption": "Fig. 2 Typical flow and pressure pattern in a centered annular seal", "texts": [ " The friction coefficients fS and fR depend on the local flow regime are computed depending on the local flow regime: Blasius law [1] for the laminar regime (Re< 1000) and Colebrook law [13] for the turbulent one (Re> 2000). Transition (1000 < Re< 2000) is dealt with by using an interpolation polynomial [14]. An annular seal is generally characterized by a subsonic (almost incompressible flow) in the inlet section that accelerates toward the exit section. In the exit section the flow regime can be either subsonic or sonic.1 The flow and pressure patterns in the annular seal are depicted in Fig. 2. The associated boundary conditions are given in Table 2. The case of a sonic outlet was thoroughly discussed in Refs. [3] and [4] and will not be discussed in the present work. For the inlet section, the flow variables (P, Vt, Vn or T) are specified outside the thin film domain. User defined coefficients (pressure loss and prerotation, U=RX) must be introduced to take into account local inertia effects that cannot be captured by the thin film equations. For example, in the inlet section the inertia effects are included in the bulk flow equations by using either the compressible or the incompressible generalized Bernoulli equation (if the Mach number of the flow is lower than 0.3). Psupply Pinlet \u00bc 1\u00fe j 1 2 1\u00fe n\u00f0 \u00deM2 inlet j j 1 (4) Psupply \u00bc Pinlet \u00fe 1\u00fe n\u00f0 \u00deq W2 inlet 2 (5) The bulk flow equations with the above mentioned boundary conditions are solved by using a finite volume method adapted for inertia dominated compressible thin film flows. Numerical details can be found in Refs. [15] and [16]. A typical flow and pressure pattern in a centered annular seal is depicted in Fig. 2. The First Order Dynamic Model. The dynamic case is tackled in the frame of the small perturbation assumption. Following this approach, the pressure (and all other field variables and their integrals) are decomposed as zero and first order variables: P z; h; t\u00f0 \u00de \u00bc P0 z; h\u00f0 \u00de \u00fe P1 z; h; t\u00f0 \u00de P1 z; h; t\u00f0 \u00dej j P0 z; h\u00f0 \u00de (6) P1 z; h; t\u00f0 \u00de \u00bc < P1c z; h\u00f0 \u00de \u00fe jP1s z; h\u00f0 \u00de\u00bd e jxt j \u00bc ffiffiffiffiffiffi 1 p (7) The resulting first order bulk flow equations have the same form as Eq. (1) with only different source terms and are Table 1 Source terms of the bulk flow equations Equation SU Continuity 1 Axial momentum h@P=@z\u00fe sSz \u00fe sRz Circumferential momentum h@P=R@h\u00fe sSh \u00fe sRh Energy (total enthalpy) q00s \u00fe q00R \u00fe RXsRh \u00fe h@P=@t Table 2 Boundary conditions for the compressible flow in an annular seal Type of boundary conditions Imposed/extrapolated flow variables Subsonic inlet (P or Vn), (Vt or Vt=Vn) and T Supersonic inlet (P and Vn) and (Vt or Vt=Vn) and T Subsonic outlet P or Vn (Vt and T) are extrapolated Supersonic outlet P, Vt, Vn, and T are extrapolated1A supersonic exit is a rare case; it could be encountered only when seal is very long and has convergent-divergent axial clearance variation", " (1) t \u00bc time (s) T \u00bc bulk temperature (K) V \u00bc total velocity (m/s) W, U \u00bc axial, circumf. bulk velocities (m/s) z, x \u00bc Rh axial and circumf. direction (m) e \u00bc e/R relative eccentricity j \u00bc ratio of specific heats n \u00bc inlet pressure loss coefficient q \u00bc density (kg/m3) s \u00bc wall shear stress (Pa) U \u00bc generic variable (1, W, U or it) x \u00bc excitation speed (rad/s) X \u00bc rotation speed (rad/s) c, s \u00bc real and imaginary part in Eq. (7) n, t \u00bc normal, tangential S, R \u00bc stator, rotor X, Y \u00bc directions depicted in Fig. 2 Figures 20\u201323 depict the numerical results obtained for the hypothetical incompressible annular seal. They were added for comparisons with Figs. 10\u201313 for the compressible annular. [1] Childs, D. W., 1993, Turbomachinery Rotordynamics, Wiley, New York. [2] Childs, D. W., and Arthur, S. P., 2013, \u201cStatic Destabilizing Behavior for Gas Annular Seals at High Eccentricity Ratios,\u201d ASME Paper No. GT2013-94201. [3] Arghir, M., Defaye, C., and Fre\u0302ne, J., 2007, \u201cThe Lomakin Effect in Anular Gas Seals Under Choked Flow Conditions,\u201d ASME J" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002808_icra40945.2020.9196650-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002808_icra40945.2020.9196650-Figure1-1.png", "caption": "Fig. 1. Proposed scenario. The mission space Q on the environment E is monitored by a team of quadcopters. Quadcopter i has a circular field of view Fi, which depends on its position pi and focal length \u03bbi. The crosshatched area in-between the quadcopters\u2019 fields of view shows a hole.", "texts": [ "00 \u00a92020 IEEE 3255 Authorized licensed use limited to: Cornell University Library. Downloaded on September 27,2020 at 11:30:23 UTC from IEEE Xplore. Restrictions apply. in Section III, under the assumption of a static graph. Section IV reconsiders the proposed scenario as a hybrid system, and extends the result in Section III into large numbers of agents with the switching of the graph. The performance of the proposed algorithm is evaluated in simulation and experiments in Section V. Section VI concludes this paper. In this paper, we consider the scenario illustrated in Fig. 1, where n quadcopters, labeled through the index set N = {1, \u00b7 \u00b7 \u00b7 , n}, equipped with cameras are distributed over 3-D Euclidean space to monitor a planar region, which is denoted as the mission space Q. The mission space is a closed and bounded convex subspace of the environment E . A density function, \u03c6 : E \u2192 R+ := [0,\u221e), encodes the importance of each point on E so that, the higher the importance, the higher the value of \u03c6 in Q, with \u03c6(q) = 0, q \u2208 E\\Q. The world coordinate frame, \u03a3w, is arranged so that its XwYw-plane is coplanar with E where the standard basis of \u03a3w is described as {ex, ey, ez}" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001129_1.4028623-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001129_1.4028623-Figure2-1.png", "caption": "Fig. 2 The discretization of the PHJM: (a) the discrete system of the PHJM and (b) c-bar unit", "texts": [ " Hence, the PM has three rotations about X-axis, Y-axis, and Z-axis, respectively, and one translation along Z-axis. Here, the spherical joint S at o represents the hip joint. In order to illustrate the moving platform\u2019s rotation movement, the PHJM topology with a general orientation is constructed in Fig. 1(b). 3.2 Position Analysis of the PHJM by FEM 3.2.1 Forward Position Analysis. When the kinematics is solved by FEM, the PHJM needs to be discretized in order to get the mechanism system composed by units and nodes according to the PHJM character (Fig. 2). In discretizing, the moving platform has four nodes according to the moving condition and constraint condition, so the moving platform is simplified to a space unit with rods 6\u201310 by equivalent transformation. Driving rod 1, rod 3, rod 4, and rod 5 are simplified to flexible rods, and the schedule functions are used to map the length variations. So the PHJM is a space system composed by 10 rods. Every rod has two sphere joints, so c-bar unit is selected as the basic unit, as shown in Fig. 2(b). The c-bar unit is a kind of space unit, and its nodes represent revolute joints in spherical joints. The stiffness matrix of the c-bar unit is \u00bdkc e \u00bc keL 2 e \u00f0Dx\u00de2 \u00f0DxDy\u00de \u00f0DxDz\u00de \u00f0Dx\u00de2 \u00f0DxDy\u00de \u00f0DxDz\u00de \u00f0Dy\u00de2 \u00f0DyDz\u00de \u00f0DxDy\u00de \u00f0Dy\u00de2 \u00f0DyDz\u00de \u00f0Dz\u00de2 \u00f0DxDz\u00de \u00f0DyDz\u00de \u00f0Dz\u00de2 \u00f0Dx\u00de2 \u00f0DxDy\u00de \u00f0DxDz\u00de \u00f0Dy\u00de2 \u00f0DyDz\u00de \u00f0Dz\u00de2 2 66666666664 3 77777777775 (20) where Dx \u00bc \u00f0xB xA\u00de \u00bc Le sin\u00f0a\u00de cos\u00f0b\u00de Dy \u00bc \u00f0yB yA\u00de \u00bc Le cos\u00f0a\u00de cos\u00f0b\u00de Dz \u00bc \u00f0zB zA\u00de \u00bc Le sin\u00f0b\u00de 9>>= >>; Le \u00bc \u00f0Dx2 \u00fe Dy2 \u00fe Dz2\u00de 9>>>>= >>>>; Geometric stiffness matrix of the c-bar unit is \u00bdgc e\u00bc \u00f0Dx\u00de2 \u00f0DxDy\u00de \u00f0DxDz\u00de \u00f0Dx\u00de2 \u00f0DxDy\u00de \u00f0DxDz\u00de \u00f0Dy\u00de2 \u00f0DyDz\u00de \u00f0DxDy\u00de \u00f0Dy\u00de2 \u00f0DyDz\u00de \u00f0Dz\u00de2 \u00f0DxDz\u00de \u00f0DyDz\u00de \u00f0Dz\u00de2 \u00f0Dx\u00de2 \u00f0DxDy\u00de \u00f0DxDz\u00de \u00f0Dy\u00de2 \u00f0DyDz\u00de \u00f0Dz\u00de2 2 666666666664 3 777777777775 (21) According to the matrix assemble principle of FEM, the geometric stiffness matrix GPHJM and the gradient vector APHJM can be expressed as follows: GPHJS \u00bc X7 i\u00bc1 \u00bd g ie APHJS \u00bc X7 i\u00bc1 \u00bda ie 9>>>= >>>; (22) According to the geometric parameters of the PHJM and the initial nodal position, the nodal position can be calculated by Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002541_s12555-019-0637-9-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002541_s12555-019-0637-9-Figure1-1.png", "caption": "Fig. 1. (a) System configuration for microrobot actuation, (b) Block diagram of feedback control for autonomous driving microrobot.", "texts": [ " Through these experimental results of the autonomous driving control of the microrobot, we measured the four driving performances (total driving time, total driving distance, stability factor, and the average distance from the obstacle) which indicate the motility of the microrobot and analyzed the autonomous driving characteristics of the microrobot using the proposed path planning algorithms. In addition, according to the changes of the mazes, we investigated and analyzed the changes in the autonomous driving performance of the microrobot. First, the system configuration for the actuation of microrobot is shown in Fig. 1(a). The EMA system composed of 8 solenoid coils was used for the autonomous driving microrobot. Each solenoid coil has a pure iron core at the center of the coil to improve the magnetic force. Here, a power supply unit (POWERSOFT, Korea) was used to apply a desired current value to each coil. The CCD camera was mounted on the top of the 8 solenoid coils and the LED light source was installed on the bottom. Therefore, by controlling the magnetic field, it is possible to manipulate the microrobot to the desired directionand position on the 2D plane [3, 24]", " In this study, because we try to compare the autonomous driving of a microrobot according to path planning algorithms in a virtual maze, we consider that it is appropriate to use the non-perfect maze which can generate various types of the path. The microrobot may have tracking errors due to unpredictable disturbances such as friction with the 2D plane surface during its operation. To reduce these tracking errors, the feedback control system with a simple proportional controller is constructed, as shown in Fig. 1(b). From the coordinate of the microrobot obtained by the threshold method of the CCD image and the desired point, an error vector can be calculated. Then, the magnetic actuating force was generated by the 8-coil system, where the magnetic actuation force has the same direction with the error vector and has the magnitude proportional to the size of the error vector. Basically, the actuation force of the microrobot is generated by the applied currents in the EMA system, and the torque and force (F) are required for the alignment and propulsion of the microrobot in the desired direction, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure8.20-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure8.20-1.png", "caption": "Figure 8.20 Drill-point thinning", "texts": [ " Smalldiameter drills will start in the correct position with the aid of the centre dot; large-diameter drills with a long chisel edge require other means to assist in starting. The best method is to use a smaller diameter drill on the centre dot, but stop before it cuts to its full diameter. The larger drill will start in its correct position guided by the 118\u00b0 dimple produced by the smaller drill. Where the chisel edge is found to be too wide for\u00a0a particular purpose, it can be reduced by point\u00a0thinning, Fig. 8.20. This can be done using the edge of a well-dressed grinding wheel, but it is\u00a0perhaps better left to a more experienced person. Although straight flute taps are the most commonly used type, other types of machine taps are available for use with high-production machines. Spiral point taps, often referred to as gun-nosed taps, have an angle ground on the inside of the flute, to push the metal cuttings (or swarf) forwards down the hole ahead of the cutting edges and are recommended for through holes" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002865_tro.2020.3031885-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002865_tro.2020.3031885-Figure1-1.png", "caption": "Fig. 1. (a) Illustration of corneal transplantation surgery using a proposed corneal suturing robot. (b) Ideal corneal suture has a suture length of 1.2 mm and a suture depth of 0.55 mm [2]. (c) If the suture length or suture depth are different between the recipient side cornea and the corneal graft, wound inversion can occur. (d) If the suture depth is too low, wound gaping may occur. (e) If it is too large, microbial invasion from the external environment can occur.", "texts": [ " Color versions of one or more of the figures in this article are available online at https://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TRO.2020.3031885 of replacement is as follows. First, the diseased cornea is cut into a circular shape and removed from the patient\u2019s eye. Next, the corneal graft is also cut into a circular shape that is similar to the patient\u2019s corneal button. Finally, the patient\u2019s eye and the corneal graft are connected using 16 sutures around the corneal edge, as shown in Fig. 1(a) [2]. During PKP, the shape of every suture should be precisely controlled to reduce visual disorders and postoperative complications. Corneal suturing differs from other suturing techniques because the suture shape considerably affects eye vision. To prevent visual disorders, there are three requirements: (R1) high suture shape uniformity, (R2) high suture shape accuracy, and (R3) avoidance of touching endothelium, the innermost layer of the cornea. (R1) Uniformity of the suture shape affects the postoperative astigmatism level", " As a consequence, extreme astigmatism can occur, which cannot be corrected by glasses or lenses and consequently requires resurgery [2], [3]. 1552-3098 \u00a9 2020 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See https://www.ieee.org/publications/rights/index.html for more information. Authorized licensed use limited to: Carleton University. Downloaded on June 04,2021 at 20:09:41 UTC from IEEE Xplore. Restrictions apply. (R2) Accuracy of the suture shape affects wound construction between the patient and the grafted cornea [Fig. 1(b)\u2013(e)]. The ideal suture shape has a suture length of approximately 1.2 mm and a depth of approximately 90% of the corneal thickness, which is approximately 550\u03bcm. However, the cornea diameter and thickness varies from person to person. Therefore, the appropriate length and depth depend on the patient. If the suture shape is not accurate and inappropriate suture depth and length are generated, various wound construction failures can occur. A short suture length can lead to wound leakage. Alternatively, a long suture length can obstruct a patient\u2019s vision" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002274_j.mechmachtheory.2019.103748-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002274_j.mechmachtheory.2019.103748-Figure2-1.png", "caption": "Fig. 2. Force Q applied far away from the centre of the contact area: (a) neglecting clearance between slider and guide, (b) considering clearance between slider and guide.", "texts": [ " 1 are valid until the line of action of the external force Q is moved beyond the contact area. When the line of action of the external force Q reaches the edge of the contact area the effective contact area becomes a single point. Under the rigid body assumption, this causes no significant impact on the friction force model, unless this model is dependent on the contact area. Beyond the edge of the contact area, however, the change in the location of the external force Q line of action has a more severe impact upon the friction force model. In Fig. 2 a, the line of action of the external force Q passes by the right side of the contact area edge. If the clearance between the guide and slider is neglected, as when a small amount of flexibility is allowed, the normal reaction is distributed between two areas of the guide on the upper and lower sides. Generally, it is assumed that the force distributions obey a triangular law and therefore the locations of the normal reactions Q N \u2032 and Q N \u2032\u2032 are known. The magnitude of the normal reactions Q N \u2032 and Q N \u2032\u2032 can be computed in terms of the external force Q , the slider length, and the moment produced by the external force Q at the slider centre", " Once the normal reactions Q N \u2032 and Q N \u2032\u2032 are known, the friction forces Q F \u2032 and Q F \u2032\u2032 can be computed based on the same friction model. One advantage of the network method used in this work is that the actions transmitted throughout the system are treated in the screw form, namely a force vector and a moment vector produced by this force at the origin. This moment vector can be computed at any point of interest by a change of coordinates. If the clearance between the guide and slider is not neglected, the slider is expected to suffer a small rotation and two contact points are established as shown in Fig. 2 b. Assuming that the clearance is small, so that the only appreciable effect produced by the rotation is the establishment of the contact points, the only noticeable difference between the situations described in Fig. 2 a and b is the location of the normal reactions Q N \u2032 and Q N \u2032\u2032 . If, for example, Coulomb friction is used, a simple scaling of the coefficient of friction by 3/2 can make the results obtained from the two approaches compatible [19] . Two modes of operation have been identified. In the first mode, illustrated in Fig. 1 , there is only one normal reaction sharing the same line of action with the external force. The second mode is illustrated in Fig. 2 . In this mode, there are two normal reactions, for which the locations of the lines of action are independent of the external force location, but the magnitudes are dependent on the external force location and on the slider width. The discrimination between these two modes is based on the external force location. Sometimes, this distinction can be made a priori . But in most cases, the possibility of two modes can make the modelling process unnecessarily complicated. When possible, simplification should be carried out to avoid these complications", " If the equivalent force and torque are known, a decision regarding the operation mode can be made and, consequently, the correct normal reaction applied. However, in a network problem, these equivalent actions are dependent on the friction force which, in turn, is dependent on the respective normal reaction. To solve this dilemma, the slider can be split up into two small parallel sliders as shown in Fig. 4 . Each slider is capable of transmitting one independent force. These forces correspond to the normal reactions of Fig. 2 or are equivalent to the single normal reaction of Fig. 1 . There is no need to determine the mode of operation with the split slider. Notwithstanding, the vertical location of the friction force is dependent on the direction of the respective normal reaction (compare Q F \u2032 in Fig. 4 a and b). If the height of the slider is small, this fact can be neglected, and in other case, the location of the friction forces is dependent on the sign of the corresponding normal reaction magnitude, i.e., the sign of the force transmitted by the coupling" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002520_j.matpr.2020.05.438-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002520_j.matpr.2020.05.438-Figure5-1.png", "caption": "Fig. 5. Blades manufactured by additive manufacturing.", "texts": [ " Thus, all the parts of CAWT are produced based on additive manufacturing concept. The parts produced by FDM method requires finishing touch. The printing surface is smoothened and irregularities if any can be removed by chemical treatment with acetone vapour, file finishing or using sand paper. The specification of additive manufacturing machine based on FDM technique available is listed in Table 3. With reference to Fig. 3, all the CAWT parts produced by additive manufacturing concept are assembled together. Fig. 5 shows the assembled CAWT and its dimensional detail [27] is listed in Table 2. Fig. 6 illustrates the experimental test setup erected at HITS, Padur, Tamilnadu to analyze the performance of CAWT. The test rig consists of an open jet wind tunnel and the rotational speed of the axial fan of the open jet wind tunnel can be varied from minimum to its maximum rated speeds. The variations in its rotational speed are done by using the ABBTM make variable frequency drive. An anemometer (cup type) is utilized to record the generated wind speeds" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002010_cca.2016.7587999-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002010_cca.2016.7587999-Figure1-1.png", "caption": "Fig. 1: Quadrotor aircraft scheme. Adapted from [17].", "texts": [ " State estimation is performed by a derivative-free nonlinear Kalman Filter, as described in [15]. The remainder of the paper is organized as follows: Section II presents a description of the quadrotor helicopter model, whose flatness property is explored in Section III. The proposed flatness-based MPC control strategy for trajectory tracking is exposed in Section IV. Section V presents simulation results to validate the proposed controller. The conclusions and suggestions for further research are given in the last section. Consider the quadrotor configuration depicted in Fig. 1. It consists of a body fixed frame B and four rotors which can generate four independent thrusts fi, i = 1, 2, 3, 4. Opposite propellers (1) and (3) turn in the same clockwise direction while propellers (2) and (4) turn in both in the counterclockwise direction. By varying the rotor speeds, one can change the lift forces, provoking the motion of the system. Thus, increasing or decreasing the four rotor\u2019s speeds together will affect vertical motion (Z-axis). Longitudinal motions (X-axis) are achieved by changing the forces f1 and f3, while lateral displacements (Y-axis) are performed through the variation on forces f2 and f4" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002736_jsen.2020.3022421-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002736_jsen.2020.3022421-Figure2-1.png", "caption": "Fig. 2 (a) Schematic diagram of the soft finger bending; (b) Movement deformation and moment distribution at any segment of elastic rod", "texts": [ " Each part roughly simulates the shape of a complete human finger, where the total length of the finger is 90 mm and the length of finger skeleton is 83 mm. The material parameters of the soft finger are shown in Table \u2160. The Cosserat elastic bar theory [29-31] takes the rod as a curve composed of a series of material points. The state of each point is only related to the material of the small area in the point. We simplify the finger as an elastic rod, whose pose state is regarded as a smooth curve in space as shown in Fig. 2(b). Fig. 2(a) shows the bending force of the finger. Fd is the driving force, FN is the grasping force received by the fingertip, and Ff is the friction force. We take the finger arc length s as the parameter to analyze the bending deformation of fingers in the O-XY plane, since the fingers are deformed in the same plane. At any point in the finger skeleton, the coordinate vector of position P is r(s)=[x(s) y(s)] and \u03b8 is the angle between the tangent direction and the Y-axis at point P. As the curvature k at ds of the differential arc on the bar can be expressed as k(s)=d\u03b8/ds, the relationship between bending moment and curvature at point s can be expressed as d ( ) d m s EI s (1) The derivative of the coordinate vector r(s) concerning s at point s on the bar is expressed as T Td d ( ) d ( ) sin cos d d d r x s y s s s s (2) Fig. 2(b) shows the distribution of force and torque from \u2018a\u2019 to \u2018s\u2019 segment where the elastic rod deforms. Note that the contact force of the rod at \u2018a\u2019 and \u2018s\u2019 are n(a) and n(s), the corresponding contact moments are m(a) and m(s), f is the distributed external force, and \u03c4 is the distributed moment. According to the equilibrium state of the elastic rod, the equilibrium equations of force and torque are ( ) ( ) ( )d 0 s a n s n a f (3) ( ) ( ) ( ) ( ) ( ) ( ) [ ( ) ( ) ( )]d 0 s a m s m a r s n s r a n a r f (4) According to the actual finger model, the distributed external force defined as the finger's own gravity can be expressed as f=[0 \u03c1Ag]T, where the finger's density is \u03c1 and the finger's cross-sectional area is A", " According to (3), the contact force at the elastic bar \u2018s\u2019 is n(s)=[a1 a2-\u03c1Ags]T. Given the stress on the fingertip n(L)=[Fx Fy]T, there is T( ) [ ]x yn s F F Ags AgL (5) Considering the finger is not subject to the distributed torque, if (5) is taken into (4), we can get ( ) ( )sin cosy xm s F Ags AgL F (6) Therefore, the static model of the finger can be expressed by the differential equation as d ( ) =sin d d ( ) = cos d d ( ) d ( ) ( )sin cosy x x s s y s s s m s EI m s F Ags AgL F (7) According to Fig. 2(a), when the driving torque on the fingertip is Td, we can obtain the finger boundary conditions (0) 0, (0) 0, (0) 0, ( ) dx y m L T (8) By combining (8) with (7), we can obtain the deformed shape of the finger. The position at point s on the rod can be expressed as 0 0 0 ( ) sin d ( ) cos d ( ) ( ) ( ) d ( ) s s s y x x s s y s s F AgL x s F y s Ags x s EI (9) Authorized licensed use limited to: UNIVERSITY OF ROCHESTER. Downloaded on September 22,2020 at 07:50:59 UTC from IEEE Xplore" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001533_0142331215619972-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001533_0142331215619972-Figure1-1.png", "caption": "Figure 1. Definition of cross-track error y, and the angles x, x R and x E .", "texts": [ " The guidance law is implemented on a test vehicle, the experimental setup is described in Section 5. Different scenarios with small and large track errors and with circular and loiter missions are flown and tested. Flight results show the efficacy of at University of British Columbia Library on March 8, 2016tim.sagepub.comDownloaded from the proposed guidance logic, these are presented in Section 6. Finally, Section 7 concludes the paper. In order to set up the guidance problem, we first define three important angles x E , x R and x as shown in Figure 1. The problem we address in this paper is guiding an air vehicle from one waypoint to the next with minimum cross-track deviation in the presence of disturbances. Here y denotes the lateral (or cross-track) displacement of the vehicle. Figure 1 indicates the positive sense of the variables. Let WP1 and WP2 be two consecutive waypoints on the earth\u2019s surface on the desired path, and let x R be the angle of the line WP1-WP2 with respect to north (called the reference or desired course angle). This is computed using latitude and longitude information of the waypoints (Samar et al., 2013). Its rate of change, i.e. _x R is zero for straight paths, and can be calculated for circular paths as: _x R = V R . The course angle x is the angle of the ground velocity V with respect to north", " Lift is divided into two components, one balances the centrifugal force and the other balances the weight of the vehicle L cosf=mg, L sinf= mV 2 R \u00f03\u00de where m is the mass, g is the gravitational acceleration, V is the ground speed, and R is the radius of turn of the vehicle. For a coordinated turn we have tanf= V 2 Rg \u00f04\u00de During a steady turn V =R _x, so equation (4) takes the form tanf= V _x g \u00f05\u00de Now since _x E = _x _x R , therefore we have tanf= V ( _x E + _x R ) g \u00f06\u00de or _x E = g V tanf _x R \u00f07\u00de where _x R is input from the mission plan. Another state equation can be derived from the the component of the ground velocity vector V ! (Figure 1) that is equal to the rate of change of lateral displacement y. We can write _y=V sin x E \u00f08\u00de at University of British Columbia Library on March 8, 2016tim.sagepub.comDownloaded from Finally the third state equation represents the inner loop dynamics, i.e. _f= 1 t (fref f) \u00f09\u00de Equations (7) to (9) represent the dynamics to be considered during outer loop guidance design. x E , y and f are the state variables, and fref is the control signal that the guidance logic generates for path following. SMC (Edwards and Spurgeon, 1998; Utkin and Shi, 1999; Zinober, 1994) is considered to be an effective and robust control design technique" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001755_978-3-319-26327-4-Figure1.10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001755_978-3-319-26327-4-Figure1.10-1.png", "caption": "Fig. 1.10 Amodel of the Girona500 AUV equipped with 5 thrusters arranged in a particular layout as shown. In the conducted failure recovery experiments, one of the surge thrusters is broken", "texts": [ " The output of this module is a vector of functionality coefficients in range [0, 1], where 0 indicates a totally nonfunctional thruster, 1 represents a fully functional thruster, and for instance, 0.7 indicates a thruster with 70% efficiency. We consider the problem of using the functional thrusters to bring the vehicle safely to a station where it can be rescued, when the thruster failure reduces the mobility of the vehicle, and hence it cannot maneuver as previously prescribed. The AUV we use for our experiments is Girona500 (Ribas et al. 2012) which is used in the PANDORA project (Lane et al. 2012), see Fig. 1.10. Girona500 is a reconfigurable AUVequippedwith typical navigation sensors (e.g.DopplerVelocityLog, etc.), basic survey equipments (e.g. side scan sonar, video camera, etc.), and various thruster layouts. In the layout we selected, the AUV is equipped with 5 thrusters: 2 heave, 2 surge, and 1 sway thrusters. We frame our approach as model-based direct policy search reinforcement learning for discovering fault-tolerant control policies to overcome thruster failures in AUVs. The described approach learns on an on-board simulated model of the AUV", " J (\u03b7) is the velocity transformation matrix, MRB is the rigid body inertia matrix, MA is the hydrodynamic added mass matrix, CRB (\u03bd) is the rigid body Coriolis and centripetal matrix, CA (\u03bd) is the added mass Coriolis and centripetal matrix, D (\u03bd) is the hydrodynamic damping matrix, g(\u03b7) is the hydrostatic restoring force vector, B is the actuator configuration matrix, and the vector \u03c4 is the control input vector or command vector. In our experiments we use Girona500 (Ribas et al. 2012) which is a reconfigurable AUV equipped with typical navigation sensors (e.g. Doppler Velocity Log Sensor), survey equipments (e.g. stereo camera) and various thruster layouts. As depicted in Fig. 1.10, the selected thruster layout in this work consists of five thrusters: 2 in heave direction, 2 in surge direction, and 1 in sway direction. In order to build a model of the system for simulating the behaviors of the AUV, the hydrodynamic parameters of Girona500, are substitute in the dynamics equations of the AUV (1.6). The hydrodynamic parameters are extracted using an online identification method and are reported in Karras et al. (2013). Policy Representation In this work we consider the control input vector u as a function \u03a0(\u03c7 |\u03b8) of observation vector \u03c7 depending on a parameter vector \u03b8 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001279_s0263574714000149-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001279_s0263574714000149-Figure8-1.png", "caption": "Fig. 8. The multi-particle system (MPS) model.", "texts": [ " With these models assumed for bodies and joints, a new criterion can be formulated for the mechanism mobility as: M = S \u00b7 p \u2212 \u2211 Ci, (9) in which p is the number of particles describing the model, S the space dimension (for a particle, S = 3 for 3D space and S = 2 for 2D space) and ci is the total number of constraints. The vector of the generalized coordinates will be obtained by numerically solving the system of M + ci algebraic equations corresponding to the M driving motions and ci joint constraints. For the sample mechanism modeled as in Fig. 8, the number of particles per body was set at 2 except for bodies 3 and 5 which are each set at 3 in order to allow the definition of the joints. The total number of mobile particles is thus p = 12 (A1, B1, B2, C2, C3, D3, E3, E4, F4, F5, G5, H5), i.e. S\u00b7p = 2\u00d712 = 24 generalized coordinates\u2014two Cartesian coordinates for each particle. http://journals.cambridge.org Downloaded: 18 Mar 2015 IP address: 128.233.210.97 As constraints, we have 9 rigid body constant distances (AB, BC, CD, DE, CE, EF, FG, FH, GH) and 14 constraints for the 7 joints, yielding ci = 23" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000367_ecc.2019.8796032-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000367_ecc.2019.8796032-Figure1-1.png", "caption": "Fig. 1: Representation of the system and its main variables.", "texts": [ " In Section III we illustrate the on-line reference generator we developed, discussing in detail the optimization problem we solve within the NMPC implementation. In Section IV the low level controller is introduced. In Section V we report the numerical results obtained and we show that the low-level controller improves the performance when external disturbances affect the system. Finally, in Section VI we gather our conclusions. Consider an unmanned aerial vehicle (UAV) tethered to the ground through a taut cable with fixed length l (see Figure 1). The robot body (B) has mass mB \u2208R>0 and inertia JB \u2208R>0, whereas the cable mass and inertia are neglected. Let us denote the world frame by FW , with origin OW and axes {xW ,yW ,zW} where zW is opposite to the gravity vector. Then let us introduce the frame FB rigidly attached to the UAV, with axes {xB,yB,zB} and origin OB, set on the center of mass (CoM) of the body. Observe that the axes yB and yW are parallel to each other and both perpendicular to the plane where the robot moves; in this paper, indeed, we assume that the UAV can move only in the 2-dimensional xW ,zW - plane", " In particular, the CoM of the UAV is constrained to move on the circle centered at OW with radius l. Thus, position and orientation can be completely described by the generalized coordinates q = (\u03d5,\u03d1) where \u03d5 represents the elevation of the UAV w.r.t. the ground, while \u03d1 is the UAV attitude; in particular, the position of the UAV is described on FW as pB = [xB,yB,zB] T = [l cos\u03d5,0, l sin\u03d5]T while the angular velocity as \u03c9 B = [0, \u03d1\u0307B,0]T . We assume the UAV is endowed with two propellers, both situated at distance d from the CoM (see Figure 1). When rotating these propellers, the forces f1, f2 are generated, where fi = b\u03c92 i , being b a constant depending on physical characteristic and \u03c92 i the angular speed of propeller i. The forces f1, f2 actuate the system with the thrust fR \u2208 R and torque \u03c4R \u2208 R (both depicted in red in Figure 1) such that fR = \u2212 fRzB and \u03c4 R = \u03c4RyB. The relation between fR, \u03c4 R and f1, f2 is uniquely determined by[ fR \u03c4 R ] = [ 1 1 \u2212d d ][ f1 f2 ] (1) In the following we will assume fR, \u03c4R to be the inputs of the UAV. In addition, the extremities of the cable are anchored to a fixed point, OW , and to OB that moves on the 2D plane. The dynamic model of the system is derived using the Euler-Lagrangian formulation, computing the kinetic and potential energies K and U , the Lagrangian function L = K\u2212U , the generalized forces Q and solving d dt dL dq\u0307\u2212 dL dq = Q, where q = [\u03d5 \u03d1 ]T; the following model is thus obtained mBl\u03d5\u0308 =\u2212mBgcos\u03d5 + fR cos\u03d5 +\u03d1 JB\u03d1\u0308 = \u03c4B" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001163_s12206-013-1134-3-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001163_s12206-013-1134-3-Figure3-1.png", "caption": "Fig. 3. Tower top coordinate system.", "texts": [], "surrounding_texts": [ "suffer from much more severe operational conditions compared to small bearings used for general industrial applications; in particular, they are subjected to greater external forces in each direction, very thin lubrication film thickness during operation, and so on. Furthermore, because they are installed on the wind turbine blade or a tower top of high altitude, it is costly and time consuming to repair or replace them in the field. Therefore, a laboratory-scale test is required for ensuring the performance of the pitch and the yaw bearings before application in the field. A test rig is developed to verify the performance of the pitch and the yaw bearings, including their fatigue life and static loading capacity. The test rig can reproduce actual operational conditions such as 6 degree of freedom (DOF) dynamic loadings and rotation of bearings for both directions. The mounting interfaces of the test rig are also the same as those used in the original environment, and various sizes of bearings can be applied by using a changeable adaptor. This high reproducibility of actual loading, driving, and mounting conditions simultaneously as well as applicability to wide size ranges are distinctively advantageous characteristics compared to previous test rigs. A structural analysis and preliminary friction torque test showed the suitability of the developed rig for use in pitch and yaw bearings of 2.0-3.0 MW class wind turbines.\nKeywords: Pitch bearing; Test rig; Wind turbine; Yaw bearing ----------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------------\nThe pitch bearing and yaw bearing of a wind turbine are mechanical rotational elements that are essential to its safe and efficient operation [1]. The former enables relative rotational motion of the blade against the hub in a certain wind speed range, and the latter enables the tower to rotate against the nacelle to track the wind direction.\nBoth bearings should be capable of supporting a very heavy load of the order of several thousand kilonewtons and should have large inner and outer diameters of the order of several meters. Furthermore, they have rotational speeds of the order of a few RPM, which is quite low compared to industrial bearings having speeds of the order of several thousand RPM. The high load and low rotational speed result in a very severe lubrication environment at the rolling element contact surface, called as boundary lubrication [2]. These unique operational conditions should be accurately captured in the design stage, and the design should be verified through a laboratory-scale test. Robust design and verification tests are indispensable\nprocesses to ensure the reliability of pitch and yaw bearings. Several international standards [3, 4] and guidelines [5, 6] deal with the design of bearings, including methods for calculating the safety factor for static loadings and the endurance life for fatigue loadings. However, there exist no standards or guidelines regarding test items or procedures to ensure design robustness. Cumulative experience and know-how is required to establish reliable test procedures, for which purpose a test rig that can closely reproduce the actual operational environment of pitch and yaw bearings is first required.\nIn this study, we developed a new test rig for the pitch and yaw bearings of a wind turbine. This rig is unique compared to previous rigs in that it can apply all possible dynamic load components, including tilting moment, and faithfully reproduce actual operational environments including driving and mounting conditions. Furthermore, it can be applied to pitch and yaw bearings of various sizes and to the performance test of pitch and yaw drives. The applicability of the test rig to 2.0- 3.0 MW class wind turbines was shown through structural analysis and preliminary friction torque tests.\nOnly a few test rigs have been developed for testing large\n*Corresponding author. Tel.: +82 42 868 7994, Fax.: +82 42 868 7477 E-mail address: yjpark77@kimm.re.kr \u2020 Recommended by Editor Sung-Lim Ko \u00a9 KSME & Springer 2014", "wind turbine bearings. Among them, Astraios from Shaeffler Technologies is the best known one (Fig. 1) [7]. This rig comprises eight cylinders\u2014four radial and four axial\u2014that can be used to represent 6 degree of freedom (DOF) loads. Various types of bearings, including general slewing bearings and main bearings, can be tested at rotation speeds of 4-20 RPM through the drivetrain connection. The testable maximum outer diameter is 3.5 m. However, this rig is more suitable for testing main bearings than for testing pitch and yaw bearings; this is because the drive type used is a drivetrain that is suitable for main bearings. In contrast, pitch and yaw bearings respectively use a pitch and yaw drive comprising a motor, reduction gearbox, and pinion as a driving system. Consequently, this test rig cannot reflect the effect of the pitch and yaw drive on the bearing performance, and this may result in some discrepancies between the test results and the actual working performance.\nMTS\u2019s test rig can be used for all large wind turbine bearings including the main bearing and pitch and yaw bearings [8]. 6-DOF loadings and oscillatory motions can be simulated. Furthermore, bearings of various sizes can be applied by using changeable adaptors. However, the drive type of this rig, too, differs from that for pitch and yaw bearings, and full rotation motion is impossible if a cylinder is used as the driving unit.\nPSL\u2019s test rig for large wind turbine bearings was originally developed for testing only main bearings [9]. It can achieve a otational speed of 25-40 RPM and bear a large axial loading of up to 5,000 kN. However, it cannot be used to represent 6- DOF loadings because it does not have a radial direction loading system.\nFor wind turbine component design, the coordinate systems used include the blade coordinate system, chord coordinate system, hub coordinate system, rotor coordinate system, tower top coordinate system, and tower bottom coordinate system [5]. The blade coordinate system and tower top coordinate system are generally applied for pitch and yaw bearings, as shown in Figs. 2 and 3, respectively.\nThe loads experienced by pitch and yaw bearings under actual operational conditions can mainly be divided into extreme and fatigue loads [5]. Extreme loads are the most severe loads applied during the service life; these are static loads that are applied for a short time. Extreme loads are normally expressed as shown in Table 1, where \u03b3F denotes the safety factor for each load case, and the entire value was arbitrarily selected. Fatigue load is a repetitive load that is applied during the service life, and it is expressed as the load duration distribution (LDD) or damage equivalent load (DEL). LDD is a load spectrum that shows all load cases with their applied time. It is a static load with only a mean value, and in some cases, it may additionally also include oscillation amplitudes and oscillation speeds. DEL is an equivalent load that causes the same damage as that caused by all types of loads applied to the bearing during its service life. In general, the rain flow counting (RFC) method is used to obtain the DEL from unordered loading data, and DEL is a dynamic load that simultaneously has both an average and a range value. Furthermore, DEL includes the reference cycle, that is, the frequency of the loading as well as the S-N curve slope, to select the proper loading level according to the properties of the used material as shown in Table 2. The values in Table 2 were also arbitrarily selected.\nAs shown in Figs. 2 and 3, the loadings applied to the pitch and yaw bearings are expressed using six types of forces and moments: Fx, Fy, Fz, Mx, My, and Mz. Of these, Fz and Mz are 2-DOF axial loads, and Fx, Fy, Mx, and My are 4-DOF radial loads. Radial loads are often combined to obtain a single root mean square (RMS) value that is expressed in terms of the resulting transverse force, Fres, and resulting bending", "moment, Mres. General-use industrial bearings are designed by considering only force components because they are usually mounted on high-speed shafts that mainly perform power transmission, and they suffer from little shaft bending moment [3, 4]. However, pitch and yaw bearings in wind turbines support heavy interfacing structures that rotate at much lower speeds. Therefore, depending on the structural flexibility of the interfacing structures, these bearings may suffer large bending moments. In fact, the bending moment is known to have the greatest influence on the fatigue life of these bearings, and therefore, it should be carefully considered during their design stage. The various loading components applied to pitch and yaw bearings can be converted into a dynamic equivalent axial load, which is an axial stationary load that has the same fatigue effect as all loading components combined. The National Renewable Energy Laboratory (NREL), USA, devised an algebraic formula that could be used to calculate the dy-\nnamic equivalent axial load using bearing specs and loading information [6]. However, because bearing failure typically occurs at the interface of the maximum loaded ball and the raceway and it is impossible to decide the effect of each of many loading components on individual ball loading, it is necessary to test pitch and yaw bearings under actual load conditions to ensure their operational reliability.\nThe pitch and yaw bearings used in wind turbines have numerous bolt holes on their inner and outer ring, and they are respectively bolted to the hub and blade and the tower and nacelle during normal operation. To test these conditions, these bearings should be bolted to the corresponding frame.\nFurthermore, pitch and yaw bearings can be of internal or external teeth type, respectively shown in Figs. 4 and 5, depending on the location of the integrated gear teeth. Pitch bearings are typically of the internal teeth type, whereas yaw bearings can be of both types. Therefore, the test rig should be designed so as to be able to test both types of bearings. For internal teeth type pitch bearings, the outer ring is bolted to the hub and remains stationary, whereas the inner ring is bolted to the blade and rotates according to the wind speed. Therefore, the bearing\u2019s inner and outer rings should be divided into loaded and stationary parts to apply test loads to the loaded part while the stationary part is connected to the fixed supporting structure.\nThe test rig should be able to load and rotate the test bearing simultaneously. It is impossible to perform the two functions simultaneously using only a set of bearings because of the structural configuration of the bearings and the test rig. To solve this problem, by using two sets of the same bearings, two rings with gear teeth attached for both bearings are con-" ] }, { "image_filename": "designv11_22_0003453_s12206-021-0334-5-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003453_s12206-021-0334-5-Figure3-1.png", "caption": "Fig. 3. Experimental setup.", "texts": [ " In addition, the stress concentration occurring at the hinges could be minimized by revising the mechanism as distributed and thereby increasing flexibility and workspace. The proposed compliant serial closed loop chain mechanism consists of two servo motors to actuate the base links, two stepper motors with a belt driving system that allows the displacement of the servo motor brackets using rail and carts purchased from Misumi, Arduino Due to control the servo motor using Matlab and Simulink, solderless breadboard to supply power and connect the servos to the Arduino, and counter weights to apply downward force on the pen as the experimental setup is shown in Fig. 3. The counterweights attached to the pen are used to darken the ink lines drawn on the board. Mechanism can either be driven by servo motors or slider to actuate the base links. If servo motors are used, then the base links follow a pure rotation and if step motors are utilized then the base links slide in the horizontal direction. Also, this experimental setup allows to actuate both servo and step motors so that the tip of the mechanism can follow parallel trajectories. Since mechanism is monolithically designed and the relative motion between each links takes place through the deflection of flexure hinges, it\u2019s expected that the hinges will be subjected to high stress when the motors are actuated" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000384_metroi4.2019.8792915-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000384_metroi4.2019.8792915-Figure5-1.png", "caption": "Figure 5. The IoT surface control system schematic diagram", "texts": [], "surrounding_texts": [ "The wheel slip velocity ( ) is defined as the difference between the longitudinal velocities of the wheel hub and the tire tread: = ( ) ( ) \u2219 (1) ( ): is the wheel hub longitudinal velocity ( ): is the wheel angular velocity : is the wheel radius For the slip factor: = ( ) ( ) \u2219( ) (2) The optimum slip factor is comprised in: 0.05 0.15 (3) The longitudinal force exerted on the tire at the contact point is expressed as: = ( ; ) where is the vertical load on tire The Pacejka (a.k.a.: Magic Formula) tire model is a specific form for the tire characteristic function, characterized by four dimensionless coefficients, B, C, D, and E, or stiffness, shape, peak, and curvature. Is given by = ( ( ) ) (4) The values of the constants in three typical road conditions are reported in the following Table I: Therefore, the traction curves, in 3 road conditions (tarmac dry and wet, the presence of snow), are given by the following diagram:" ] }, { "image_filename": "designv11_22_0001620_s12206-015-0908-1-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001620_s12206-015-0908-1-Figure6-1.png", "caption": "Fig. 6. Static stress and strain state under full throttle cycle.", "texts": [], "surrounding_texts": [ "During the operation of an aero-engine, the turbine disc is subjected to a mixed load: centrifugal forces, thermal stresses, aerodynamic forces and vibratory stresses. Of course, high speed results in large centrifugal forces and high thermal gradients result in thermal stresses. Among them, the aerodynamic forces and vibratory stresses have little effect on the static strength of the turbine disc. Therefore, when analyzing the turbine disc with finite element method, the centrifugal forces and thermal stresses are the main consideration. The speed spectrum of the turbine disc is determined by the flight mission, and it consists of three parts [11]: low frequency cycle, full throttle cycle and cruise cycle. Any speed spectrum can be considered as a combination of these three basic cycles. The speed spectrum of the turbine disc is shown in Table 3. The temperature spectrum is derived based on the measurement data. In this study, the temperature spectrum of the turbine disc was loaded on the three-dimensional model by ANSYS parametric design language. For each basic cycle mentioned above, there are 100000 temperature data points of the turbine disc. Table 4 shows part of the temperature data points under full throttle cycle, where X, Y and Z represent the coordinate value of a point of the three-dimensional model." ] }, { "image_filename": "designv11_22_0001648_j.ijleo.2016.02.077-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001648_j.ijleo.2016.02.077-Figure1-1.png", "caption": "Fig. 1. Typical planar engagement geometry.", "texts": [], "surrounding_texts": [ "O S l\nD S\na\nA R A\nK S m F M T\n1\nr c i c a t m e z\nd a u a ( [ t b o a t\nh 0\nOptik 127 (2016) 5359\u20135364\nContents lists available at ScienceDirect\nOptik\njo ur nal homepage: www.elsev ier .de / i j leo\nriginal research article\ntochastic fast smooth second-order sliding modes terminal guidance aw design\nan-xu Zhang, Yang-wang Fang, Peng-fei Yang \u2217, Yang Xu chool of Aeronautics and Astronautics Engineering, Air Force Engineering University, Xi\u2019an 710038, China\nr t i c l e i n f o\nrticle history: eceived 22 January 2016 ccepted 29 February 2016\neywords:\na b s t r a c t\nAiming at handling the track imprecision caused by inertial lag, model uncertainties and atmospheric environment disturbances, as well as stochastic noises, a terminal guidance law based on stochastic fast smooth second-order sliding modes control theory is proposed. This paper considers targets performing evasive maneuvers and develops a high-order sliding mode observer. A concept of finite-time mean-\ntochastic fast smooth second-order sliding ode inite-time ean-square practical convergence erminal missile guidance law\nsquare practical convergence, considering the non-equilibrium additive noise of the guidance system, is presented. And according to this concept, the finite-time convergent guidance law is deduced. The feasibility of the new guidance law is exemplified through computer simulations and the guidance performance is compared with augmented proportional navigation guidance law, sliding mode guidance law and nonsingular terminal sliding mode guidance law.\n\u00a9 2016 Elsevier GmbH. All rights reserved.\n. Introduction\nAn important criterion of a homing missile is the tracking accuacy, which is closely related to guidance, navigation, and control rucially [1]. To achieve higher guidance precision against everncreasing performance targets, the line-of-sight (LOS) rate needs onverge to zero fast, which makes the terminal trajectory straight nd the normal acceleration of the missile small. However, due to he missile seeker detection lag, guidance update rate limitation,\nissile rudder inertial delay, model uncertainties and atmospheric nvironment disturbances [2], the LOS rate cannot be converge to ero within a short time.\nA series of sliding mode control (SMC) algorithms has being evoted to design the homing missile guidance law due to its dvantages of handling bounded uncertainties, disturbances and nmolded dynamics [3]. The SMC guidance law achieved smaller cceleration ratio compared to traditional proportional navigation PN) and augmented proportional navigation (APN) guidance laws 4,5]. However, classical SMC cannot ensure the LOS rate converge o zero in finite-time [6]. So Zhou presented a new guidance law ased on SMC that can guarantee the LOS rate converge to zero r its small neighborhood in finite-time [7]. Whereas, the guidnce law designed in the framework of first-order SMC requires he system relative degree equal to 1 with respect to the sliding\n\u2217 Corresponding author. Tel.: +86 029 84787589. E-mail address: pfyang1988@126.com (P.-f. Yang).\nttp://dx.doi.org/10.1016/j.ijleo.2016.02.077 030-4026/\u00a9 2016 Elsevier GmbH. All rights reserved.\nvariable and the controller yields a heavy chattering [8\u201310]. And in Ref. [8\u201310], to deal with the intrinsic difficulties of classical SMC, the high-order sliding mode (HOSM) controllers are presented. A new smooth second-order sliding mode (SSOSM) control driven by uncertain sufficiently smooth disturbances is proposed and proved by Shtessel [11,12]. The main limitations of this guidance law are the simplification of the state noise and the dependences of the perfect knowledge of the range to target and the range rate, which is usually hard to get an accuracy value. Another limitation of this method is the specific requirement of the target normal acceleration.\nIn order to solve the defects in previous research, a novel stochastic fast smooth second-order sliding mode (SFS-SOSM) method with a finite-time convergence in the presence of evasive target maneuvers, uncertainties, disturbances and stochastic noise, is proposed in this paper. A new concept of finite-time mean-square practical (FTMSP) stability is introduced to investigate the finite-time convergence of the stochastic sliding surface and the FTMSP convergence of SFS-SOSM control is proved by It\u00f4\u2019s formula.\nThe paper is organized as follows: Section 2 states the missiletarget engagement kinematics. The SFS-SOSM control algorithm is derived and its FTMSP convergence is proved in Section 3. In Section 4, a smooth guidance law based on SFS-SOSM is presented and its performance is verified via computer simulations compared with augmented proportional navigation guidance law (APN), sliding mode guidance law (SMG) and nonsingular terminal sliding mode guidance law (NT-SMG) in Section 5.", "5360 D.-x. Zhang et al. / Optik 12\n2\n2\nt F\nd{\nw t a\ns e\nq\nw t d f\n\u03c9\nw z t c\n\u03c9\n2\ni n i r r m\nw\n. Planar engagement model and intercept strategy\n.1. Problem formulation\nConsider the planar homing case that the missile moves within he vertical plane, a typical engagement scenario is presented in ig. 1.\nThe planar missile-target engagement kinematics can be easily erived as\nr\u0307 = VT cos(q \u2212 T ) \u2212 VM cos(q \u2212 M) rq\u0307 = \u2212VT sin(q \u2212 T ) + VM sin(q \u2212 M) (1)\nhere q is the LOS angle, r is the range along LOS, VT and VM are arget velocity and missile velocity, T and M are target aspect ngle and missile lead angle.\nThe eq. (1) can be reduced by differentiating both sides of the econd equation with respect to time and substituting the first quation into it leads to the following equation\n\u00a8 = \u22122r\u0307 r q\u0307 \u2212 1 r aM + 1 r aT (2)\nhere aM is missile normal acceleration as a control input, aT is arget normal acceleration that is considered as unknown bounded isturbance. Denote \u03c9 = q\u0307 as the LOS rate, Eq. (2) becomes the ollowing equation\n\u02d9 (t) = \u22122r\u0307(t) r(t) \u03c9(t) \u2212 1 r(t) aM + 1 r(t) aT (3)\nhere the starting time of the guidance process is taken to be ero, \u03c9 is an uncertain sufficiently smooth function. Assume that he state noise (t) is a zero-mean white Gaussian process with ovariance Q(t), Eq. (3) can be rewritten as\n\u02d9 (t) = \u22122r\u0307(t) r(t) \u03c9(t) \u2212 1 r(t) aM + 1 r(t) aT + (t) (4)\n.2. Intercept strategy\nIt is well known that in the space interception where a missile s intercepting a target with maneuverability, the time of termial guidance is only several seconds such that the guidance law s required to ensure finite time convergence of the LOS angular ate [7,13]. To ensure finite time convergence of the LOS angular ate, the guidance law is derived to stabilize the system (4) on the anifold \u222b\n= \u03c9(t) + t\n0\n\u03c9( )d (5)\nhere = const . >0.\n7 (2016) 5359\u20135364\nDifferentiating both sides of Eq. (5) with respect to time, we arrive at the following equation \u0307 = \u03c9\u0307 + \u03c9 = ( \u2212 2r\u0307\nr\n) \u03c9 \u2212 1\nr aM + 1 r aT + (6)\nThe guidance command can be obtained by employ SFSSOSM control, which is derived and analyzed in the next section.\n3. Stochastic fast smooth second-order sliding mode control\nIt is obvious that system (6) is driven by additive noise, meaning that the equation doesn\u2019t have any equilibrium [15]. Instead of convergence to the origin, the more reasonable way is to stabilize to a small neighborhood of zero in finite time [7,14]. Practical stability, proposed by La Salle and Lefschetz [16], is motivated by the fact that the state of a physical system may be mathematically unstable, but it operates sufficiently near the desired state. With the aid of this concept, a new concept of FTMSP convergence is introduced first, and then the SFS-SOSM control is derived and its FTMSP convergence is proved in this section.\n3.1. Finite-time mean-square practical stability\nThe definition of practical stability given in [14] is extended to a stochastic nonlinear system as follows.\nDefinition (FTMSP convergence): Denote x(t) the solution process of system (4) under the initial condition x(t0) = x0. The sliding surface = (x(t)) = 0 is called finite-time mean-square practical (FTMSP) convergent if, given real number pair \u0131, \u03b5 > 0 satisfying certain conditions, there exists a finite setting time T \u2265 0, which is\ndependent on x0, such that E \u2225\u2225 (t0) \u2225\u22252 \u2264 \u0131 implies E \u2225\u2225 (t) \u2225\u22252 \u2264 \u03b5 for any t \u2212 t0 > T.\nIt follows from the above definition that the E \u2225\u2225 (t) \u2225\u22252 is sufficiently close to zero in finite-time if the sliding surface is FTMSP convergent.\n3.2. Prescribed sliding variable dynamics\nOn the ground of the smooth second-order sliding mode (SOSM) control proposed by Shtessel [11], an extended stochastic fast smooth SOSM (SFS-SOSM) control can be deprived and the dynamics of the sliding variable is designed to have the following form:{\n\u03071 = \u2212k1 \u2223 1 \u2223(m\u22121)/msgn( 1) \u2212 k2 1 \u2212 k3 \u2223 2 \u2223 sgn( 1) + (\u03c9 \u2212 z1) +\n\u03072 = \u2212k4 \u2223 1 \u2223(m\u22122)/msgn( 2) \u2212 k5 2\n(7)\nwhere 1 = , m = const . >2, ki = const . >0 (i = 1, 2, 3, 4, 5), is the noise signal mentioned in Eq. (4).\nLet = [ 1, 2]T, then Eq. (7) is a stochastic system with respect to state and can be represented as\nd = f ( )dt + gdW(t) (8)\nwhere W(t) is a 1-dimensional standard Brownian motion and f, g are\nf ( ) = \u2212k1 \u2223\u2223 1 \u2223\u2223(m\u22121)/m sgn( 1) \u2212 k2 1 \u2212 k3 \u2223\u2223 2 \u2223\u2223 sgn( 1)\n\u2212k4 \u2223\u2223 1 \u2223\u2223(m\u22122/m) sgn( 2) \u2212 k5 2\ng = [\u221a Q\n0\n] (9)\nSystem (8) is a stochastic nonlinear system with additive noise. Hereafter, FTMSP convergence is employed to analyze the finitetime convergence of system (8).", "tik 12\n3\nd\nT s\nw\n\u03b5\n\u0131\nw c\nP\nV\nv c a f\nt\nV\nw\nm b\nw a i e\nE\nL\nw t\nD.-x. Zhang et al. / Op\n.3. Finite time convergence of SFS-SOSM\nBased on the definition proposed in Section 3.1, sufficient conitions for the FTMSP convergence of system (8) can be proved.\nheorem 1. Consider the stochastic nonlinear system (8) and contruct a matrix as follows\n= m\n2m + 1\n[ k5 0\n0 k2\n] (10)\nhere m > 2, ki > 0 (i = 2, 5) are aforementioned parameters. If there exists real numbers \u03b5 and \u0131 satisfying\n\u2265 max( ) min( ) Q 2k2 (11)\n> Q\n2k2 (12)\nhere Q is the covariance of noise . Then the system (8) is FTMSP onvergence with respect to (\u0131, \u03b5).\nroof. For system (8), define Lyapunov functional candidate as\n( ) = 1 2 (k5\n\u2223\u2223 1 \u2223\u22232 + k2 \u2223\u2223 2 \u2223\u22232 ) (13)\nSince V( ) is continuous but not differentiable, a nonsmooth ersion of Lyapunov\u2019s theory is required, which shows that one an just consider the points where V( ) is differentiable [17]. This rgument is valid in all the proofs of the present paper, so that no urther discussion of these issues will be required.\nThe substitution = [ \u2223\u2223 1 \u2223\u2223 , \u2223\u2223 2 \u2223\u2223]T brings the proposed func-\nional (13) to a quadratic form\n= T (14)\nhere\n= 1 2\n[ k5 0\n0 k2\n] (15)\nNote that V is positive definite and radially unbounded since > 2, ki > 0 (i = 2, 5), the following inequalities can be obtained\nased on Rayleigh\u2013Ritz theorem min( )E( \u2225\u2225 \u2225\u22252 ) \u2264 EV \u2264 max( )E( \u2225\u2225 \u2225\u22252 ) (16)\nhere \u2225\u2225 \u2225\u22252 = \u2223\u2223 1 \u2223\u22232 + \u2223\u2223 2 \u2223\u22232 is the Euclidean norm of , min( )\nnd max( ) are minimal and maximal eigenvalues of . Considerng that and i(i = 1, 2) are stochastic process, the mathematical xpectations satisfy\n\u2225\u2225 \u2225\u22252 = E \u2223\u2223 1 \u2223\u22232 + E \u2223\u2223 2 \u2223\u22232 (17)\nBy applying It\u00f4\u2019s formula to V yields\nV = \u2202V( , t) \u2202t\n+ (\n\u2202V( , t) \u2202\n)T\nf + 1 2 trace\n( gT \u22022\nV( , t) \u2202 2 g\n) (18)\nhere L is the infinitesimal generator. Eq. (18) can be expanded and he following inequalities hold\n7 (2016) 5359\u20135364 5361\nLV = \u2202V\n\u2202t +\n( \u2202V\n\u2202\n)T\nf + 1 2 trace\n( gT \u22022 V\n\u2202 2 g\n)\n= [k5 \u2223\u2223 1 \u2223\u2223 sgn( 1), k2 \u2223\u2223 2 \u2223\u2223 sgn( 2)][\u2212k1 \u2223\u2223 1 \u2223\u2223(m\u22121)/m sgn( 1)\n\u2212 k2 1 \u2212 k3 \u2223\u2223 2 \u2223\u2223 sgn( 1) \u2212 k4 \u2223\u2223 1 \u2223\u2223(m\u22122)/m sgn( 2) \u2212 k5 2]\n+ 1 2 trace\n( gT \u22022 V\n\u2202 2 g\n) = \u2212k1k5 \u2223\u2223 1 \u2223\u2223(2m\u22121)/m \u2212 k2k5 \u2223\u2223 1 \u2223\u22232\n\u2212 k3k5 \u2223\u2223 1 \u2223\u2223 \u2223\u2223 2 \u2223\u2223 \u2212 k2k4 \u2223\u2223 1 \u2223\u2223(m\u22122)/m \u2223\u2223 2 \u2223\u2223 \u2212 k2k5 \u2223\u2223 2 \u2223\u22232\n+ 1 2 trace\n( gT \u22022 V\n\u2202 2 g\n) \u2264 \u2212k2k5( \u2223\u2223 1 \u2223\u22232 + \u2223\u2223 2 \u2223\u22232 )\n+ 1 2 trace\n( gT \u22022 V\n\u2202 2 g\n) = \u2212k2k5 \u2225\u2225 \u2225\u22252\n+ 1 2 trace\n( gT \u22022 V\n\u2202 2 g\n) (19)\nDenote the second term of Eq. (18) as V2 and the inequality below can be deduced according to the properties of trace of matrix.\nV2 = 1 2 trace\n( gT \u22022 V\n\u2202 2 g\n) = 1\n2 gT Vg\n= 1 2\n[\u221a Q 0 ] \u23a1 \u23a2\u23a2\u23a3 \u22022 V \u2202 2 1\n\u22022 V\n\u2202 1\u2202 2\n\u22022 V\n\u2202 2\u2202 1\n\u22022 V \u2202 2 2\n\u23a4 \u23a5\u23a5\u23a6 [\u221a Q\n0\n]\n= 1 2\n\u221a Q\n[ \u22022 V\n\u2202 2 1\n\u22022 V\n\u2202 2\u2202 1\n] [\u221a Q\n0\n] = 1\n2 Q\n\u22022 V \u2202 2 1\n= 1 2 Q \u2202\n\u2202 1\n( \u2202V\n\u2202 1\n) = 1\n2 Q\nd\nd 1 (k5\n\u2223\u2223 1 \u2223\u2223 sgn( 1)) = 1 2 Qk5 (20)\nSubstitute Eq. (20) into Eq. (19) to get\nLV \u2264 \u2212k2k5 \u2225\u2225 \u2225\u22252 + k5\n2 Q (21)\nWith the aid of It\u00f4\u2019s theorem [18] (EV) \u2019 = E(LV) and Eq. (16), it follows that (EV)\u2032 = E(LV) = \u2212k2k5E( \u2225\u2225 \u2225\u22252 ) + k5\n2 Q \u2264 \u2212k2k5\nEV\nmax( )\n+ k5\n2 Q = \u2212 1EV + 2 (22)\nwhere\n1 = k2k5 max( ) , 2 = k5 2 Q (23)\nIn view that m > 2, ki > 0 (i = 1, 2, . . ., 5), we see that 1, 2 > 0. Since the solution of the differential equation\n\u03d5\u0307 = \u2212 1\u03d5 + 2, \u03d5(t0) = \u03d50 \u2265 0 (24)\nis given by \u03d5(t) = ( \u03d50 \u2212 2\n1\n) e\u2212 1(t\u2212t0) + 2\n1 (25)" ] }, { "image_filename": "designv11_22_0003026_j.egyr.2020.11.133-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003026_j.egyr.2020.11.133-Figure1-1.png", "caption": "Fig. 1. The double-chamber MFC.", "texts": [ " The anodes were set up vertically to the surface of the solution to prevent waste matters from attaching to the surface of the anodes. WT and \u2206fliC were separately inoculated in two microtubes containing 2 ml LB medium for 48 h. Subsequently, RCA-01 anodes with the size of 1 \u00d7 2 cm2 were respectively put in these microtubes for 24 h. OD600 nm of the inoculated solutions was measured when immersing the anodes and after immersing for 24 h. Also, the biofilms on the surface of the anodes were observed by a scanning electron microscope (SEM). N B e b e 3 w w Fig. 1 shows the design of the double-chamber MFC used in this study. The anode chamber was filled with 6 ml of a rich-nutrient medium composed of 10 g/l glucose, 5 g/l yeast extract, 10 g/l NaHCO3, 8.5 g/l NaH2PO4, and aOH to adjust pH = 7.0 [10]. The cathode chamber was filled with 6 ml of potassium ferricyanide (50 mM). oth the anodic and cathodic materials were RCA-01. The voltage between the two electrodes was recorded by a data acquisition system (NI USB-6210). The xperimental conditions were open-air at room temperature" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003487_00325899.2021.1901398-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003487_00325899.2021.1901398-Figure10-1.png", "caption": "Figure 10. Evolution from CAD model without (a) and with capsule (b) to optimised capsule geometry (c) and near-net-shape geometry (with capsule) after HIP simulation (d).", "texts": [ " The coordinates of all nodes are adjusted according to method described in Section 2.6. After the second simulation step, the shape is closer to the target geometry. After five iterations, the capsule geometry after simulation (iteration 5, after HIP) is virtually identical to the target geometry. The capsule geometry of this last step before HIP (iteration 5, before HIP) could be used to produce the capsule by LPBF. The second geometry to be optimised was the capsule for a screw extruder segment. The procedure of the optimisation process is shown in Figure 10. Starting with target geometry of the final part after HIP, generated from the CAD-file in Figure 10(a), a FE model was developed by automated discretisation in Figure 10(b). From this FE model, the optimised capsule geometry Figure 10(c) was identified after five iterations. This geometry file was used to produce capsule by the LPBF process. Figure 10(d) shows the predicted shape after densification of the optimised model by HIP.Figure 9. Iterations for net-shape optimisation of LPBF capsule (iterations no. 3 and 4 not shown) All capsules were subjected to the same HIP parameters. The applied HIP cycle is visualised in Figure 11. Small test capsules before and after HIP are shown in Figure 12. LPBF-built capsules have been joined to a filling tube which was closed by forging after powder filling and evacuation (Figure 12 (a)). Figure 12(b) reveals the same capsules in their deformed state after HIP", " The deviation colour representation from red over neutral white to blue [31] shows the difference between simulation and measurement. Adjacent to the colour bar, a histogram is shown, which give the distribution of the deviations. The histograms in Figures 17 and 18 are normalised to the same height and width. The filling tubes are not part of the comparison; therefore, they have not been coloured but shown in dark grey. For the cylindrical capsule shown in Figure 9, rotational symmetry was assumed in the FE model. Figure 10 depicts the example of a screw extruder element where the optimisation method also works using a full three-dimensional model. For the \u2018after HIP\u2019 state, the geometry of the optimised capsule after HIP simulation was compared to the geometry of the real capsule after HIP process. A comparison of numerically determined geometry (Figures 9 and 11(d)) and produced part in the as HIP state reveals at which areas and to what extent the HIP simulation deviates from the experiment. As can be seen in Figures 17 and 18, the simulation can quantitatively predict the shape change of a capsule during HIP" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000207_2019010-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000207_2019010-Figure6-1.png", "caption": "Fig. 6. Finite element model of bearing house.", "texts": [ " As shown in Table 2, the maximum film pressure obtained by this study agree well with those obtained by Sun. Detailed parameters of the journal bearing used in the following numerical analysis are listed in Table 3. The oil is assumed to be an incompressible Newtonian fluid, which is supplied to the bearing through the axial grooves located at upper half bushing. As the oil supply pressure is far less than the film pressure, the zero pressure can be fixed over the groove, as shown in Figure 5. The compliance matrix of the bearing house is obtained by the finite element method, which is shown in Figure 6. Grids on the bushing surface are controlled to assure the correspondence between the nodes on the bushing surface and the nodes used to calculate film pressure. The unit type is Solid185, and the constraint is applied to all nodes located at the outer surface of the bearing house. Mesh refinement analysis is performed based on the plain profile bearing with aligned journal, and the minimum film thicknesses (hmin) for different meshes are showed in Figure 7. It can be seen that the solution is converged when the mesh is 80 60" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001827_aer.2016.50-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001827_aer.2016.50-Figure1-1.png", "caption": "Figure 1. Homing engagement geometry.", "texts": [ " Considering the fact that the well-known separation principle is usually not valid for non-linear observer plus controller structure, a two-step method is utilised to prove the stability of the closed-loop guidance system. This remainder of this paper is organised as follows. In Section 2, the background and preliminaries are stated. In Section 3, the original continuous guidance law is provided, followed by the composite guidance law proposed in Section 4. Finally, the simulation results and some conclusions are offered. 2.0 BACKGROUND AND PRELIMINARIES The planar homing engagement geometry between missile and target is depicted in Fig. 1, where the subscripts M and T denote the missile and the target, \u03b3M and \u03b3T the missile and the target flight-path angle, \u03bb and r the LOS angle and the missile-target relative range, VM and VT the missile and the target velocity; aM and aT the missile and the target acceleration, which are assumed normal to their own velocities, respectively. http:/www.cambridge.org/core/terms. http://dx.doi.org/10.1017/aer.2016.50 Downloaded from http:/www.cambridge.org/core. Macquarie University, on 03 Dec 2016 at 12:28:18, subject to the Cambridge Core terms of use, available at Based on principle of kinematics, the differential equations describing the relative motion are formulated as r\u0307 = VT cos (\u03b3T \u2212 \u03bb) \u2212 VM cos (\u03b3M \u2212 \u03bb) , \u2026 (1) \u03bb\u0307 = 1 r [VT sin (\u03b3T \u2212 \u03bb) \u2212 VM sin (\u03b3M \u2212 \u03bb)] , \u2026 (2) \u03b3\u0307M = aM VM , \u2026 (3) \u03b3\u0307T = aT VT \u2026 (4) Assumption 1: The target velocity and acceleration satisfy that VT < VM, |aT | \u2264 1, \u2026 (5) where 1 > 0 denotes the upper bound of target manoeuvre" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003867_s40516-021-00150-6-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003867_s40516-021-00150-6-Figure2-1.png", "caption": "Fig. 2 Laser-arc hybrid welding with a gap in (a) butt, (b) lap-fillet, and (c) overlap configuration", "texts": [ " In contrast, arc welding may bridge a larger gap, but its power density is limited. The amalgamation of the two processes (Fig.\u00a01) overcomes the mutual limitations and offers additional advantages like preheating or post-heating either of the welding sources. The benefits of hybrid welding to improve the penetration and gravity filling of larger weld deviations in butt welds are well documented. The LAHW is studied to bridge the gap in different configurations of weld joints, namely butt, lap-fillet, and overlap (Fig.\u00a02). The gap bridgeability in butt configuration is reported more often than in the lap configuration. The gravitational force and arc force help gap filling in the butt configuration (Fig.\u00a02a). The molten metal flux is directed to the bottom of the weld pool, which heats the weld root [1]. In the lapfillet configuration, filling predominantly depends on the capillary action (Fig.\u00a02b, making it difficult to bridge the gap. The gap acts as a vent hole in the overlapping pattern (Fig.\u00a02c) to reduce the porosity [2]. In the butt configuration, when the GMAW torch follows the arc, the molten droplets move towards the joint edges, thereby filling the gap; however, the optimal parameters that work for GMAW more often are not optimal for the LAHW [3]. A seam tracking sensor and adaptive control of welding parameters can bridge a gap greater than 1\u00a0mm [4]. However, larger gaps can result in insufficient melting in the lower part of the weld, as the laser does not succeed to reach there [5]", " Deviation tolerances in hybrid laser-arc thick plate welding could be improved by additional methods such as the electromagnetic welding pool backing system [7]. The studies on the effect of the gap in the lap-fillet configuration 1 3 are reported for hot-wire laser welding [8] and laser welding [9, 10]. The gaps considered in the laser welding studies have been very narrow, such as 0.2\u00a0mm and 35\u00a0\u00b5m, providing a vent hole for reducing porosity rather than gap bridgeability. The ability to fill the gap in fillet condition for a typical lap-fillet weld, as shown in Fig.\u00a02b, is still an open question, barring very few preliminary investigations [11, 12]. The gap bridgeability is influenced by the interaction between the laser beam and the arc process. The laser stabilizes the welding arc. The molten material could flow into the gap because the capillary action to fill the gap needs further investigation. The throat thickness is the main feature that governs the strength of the weld joint. However, the gap between the plates\u2014designed or due to misalignment\u2014alters the throat thickness and affects the strength" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003664_tia.2021.3084549-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003664_tia.2021.3084549-Figure17-1.png", "caption": "Fig. 17. Photographs of a designed prototype.", "texts": [ " The temperature rise of coils when the maximum currents are applied continuously from initial temperature of 20\u00b0C is assuming that there is no heat dissipation to outside of the coils. Even in armature windings with high current density, the temperature rise for 60 seconds is \u039444.9 K. In general, the short-time rating required for a traction motor when the vehicle is running uphill is about 30 to 60 seconds. The proposed motor has enough heat capacity to drive at maximum torque for short-time rating. III. MANUFACTURE AND EXPERIMENTAL RESULTS OF PROTOTYPE We made a prototype motor based on the structure described in the previous section. Fig. 17(a) shows the exterior Fig. 12. Motor characteristics when pole arc angle of the PM \u03b8 is changed (3D-FEA result). 0 100 200 300 400 500 600 86.0 87.0 88.0 89.0 90.0 91.0 92.0 17 18 19 20 21 22 23 24 Ir o n lo ss [ W ] T o rq u e [ N m ] E ff ic ie n c y [ % ] Pole arc angle of PM \u03b8 [deg] Max. torque Efficiency Iron loss Fig. 13. Rotor as a conventional axial gap motor. Fig. 14. Maximum torque of a conventional axial gap motor and proposed motor (3D-FEA result). 0 10 20 30 40 50 60 70 80 90 100 Conventional motor Proposed motor M ax im u m t or q ue [ N m ] \uff0b89.3% \uff0d68.2% \uff0b31.3% Field weakening Field strengthening No field current Authorized licensed use limited to: Univ of Calif Santa Barbara. Downloaded on June 30,2021 at 12:28:31 UTC from IEEE Xplore. Restrictions apply. 0093-9994 (c) 2021 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. of the prototype, Fig. 17(b) shows a stator of prototype, and Fig. 17(c) shows a rotor of prototype. The field winding is placed into the inner circumference of the stators, effectively using the dead space in the motor. The terminal of the field winding is drawn out from the through hole provided in the motor case lid fixing the stator. The PM and SMC are fixed to a nonmagnetic rotor support component using an adhesive. Fig. 18 shows the no-load line-to-line voltage waveform of the prototype at base speed (1600 rpm). When the field current is changed from no field current to maximum field strengthening, the voltage amplitude increases by 94" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002301_s12221-020-9632-2-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002301_s12221-020-9632-2-Figure1-1.png", "caption": "Figure 1. (a) Stacking sequences of the fabric layers, (b) schematic of the VARTM process, and (c) cure cycle for hybrid reinforced composites.", "texts": [ " Vinyl ester (VE) of bisphenol A type resin was purchased from Polynt composites. Methyl ethyl ketone peroxide (MEKP) and cobalt naphthalate (CN) were used as curing agent and accelerator of VE resin, respectively. The hybrid composites were fabricated by Vacuum Assisted Resin Transfer Mold (VARTM) process, as explained in the literature [14]. In brief, a total of five composites were fabricated; the stacking sequences are listed in Table 2, while the schematic diagrams of the stacking sequences are shown in Figure 1. The fabrics were stacked up, and the vacuum was applied with a vacuum pressure of 0.06 MPa, and the impregnation was with VE (VE: MEKP: CN \u2013 100:1:1). The composites were cured according to the cure cycle shown in Figure 1(c). Volume Fraction and Porosity of Composites The fiber volume fraction and porosity of all specimens were measured according to ASTM D792. The specimens were cut into 10\u00d710 mm and dried in an oven for 24 h at 100 \u00b0C before weighing. The volume of the specimens was measured after they were removed from the oven. The density of the specimen was measured by the Archimedes method. The difference between the theoretical density and the measured density equals the porosity of the hybrid composite. The fiber volume fraction indicates the remaining fiber weight measured after burning the composites completely at 600 \u00b0C" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003249_tte.2021.3049466-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003249_tte.2021.3049466-Figure1-1.png", "caption": "Fig. 1. Cross section of 12-slot/ten-pole six-phase HS-RDPM.", "texts": [ " For verifying the superiority of above strategies, the opencircuit back-EMFs, torques, and fault-tolerant performances are compared between HS-RDPM with conventional FTPM and intermediate structures, which just has partial advantages of HS-RDPM. The results show that the HS-RDPM has a higher torque density than conventional FTPM. The bigger torque ripple and weak phase isolation originating from dual stators can also be improved. In addition, the loss distributions and the security of the machine under a short circuit are considered. An HS-RDPM is developed, and the experimental results agree with the simulations, verifying the analyses on performances of HS-RDPM. Fig. 1 shows the cross section of the proposed HS-RDPM with 12 slots, ten poles, and six phases. It contains three parts: outer stator, rotor, and inner stator. The windings of the same phase in different stators are connected in series. The slot area of the outer stator is bigger than that of the inner stator in consideration of the different heat-dissipation level. To protect the rotor, the PMs are embedded into the rotor frame, and several bolts with high strength have also been added to reinforce the rotor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003046_j.procir.2020.01.186-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003046_j.procir.2020.01.186-Figure1-1.png", "caption": "Fig. 1 Transient melt pool dynamics in laser drilling, P= 200W \u03c4=0.1 ms (a) after 5 (b) after 15 (c) after 20 pulses; P= 200W \u03c4=1 ms (d) after 4 (e) after 8 (f) after 10 pulses. The black contour line represents T= T_melting separating solid and molten phase.", "texts": [ "1 Melt hydrodynamics in laser micro drilling Laser drilling can be classified into two regimes, short/ultra-short pulse drilling where material is removed primarily due to vaporization and long pulse (\u03bcs-ms) drilling in which expulsion of melt creates a hole. The melt hydrodynamics in long pulse drilling is governed by the factors such as recoil pressure, surface tension, Marangoni convection and evaporative cooling. The combined effect of above parameters results in time varying flow patterns during laser heating as well as cooling (pulse off duration). Fig. 1 (a-c) illustrates different stages of lase drilling in case of P=200 W (laser intensity ~ 109 W/m2), pulse width= 0.1 ms at 5, 15 and 20 pulses respectively, pertinent time steps have been chosen to illustrate different regimes of melt dynamics. After irradiation of 5 pulses, the melt layer is thin with small melt humps near the hole edge. With increase in laser pulses, the drill depth increases albeit with multiple perturbation in the melt zone. The expulsion of melt has deposited significant amount of molten metal near the hole edge resulting in notorious burr formation in laser drilling. After the 20th pulse, the excessive melt does not fully solidifies resulting in hole blockage. Fig 1 (d-e), drilling phases for P=200 W (laser intensity ~ 109 W/m2), pulse width=1 ms at 4, 8 and 10 pulses showcase melt expulsion, melt shadowing and hole blockage. The formation of hole blockage can be attributed to decrease in melt expulsion efficiency with increase in depth. The effects associated with melt flow become more pronounced with increase in pulse width (0.1 ms-1 ms), as laser interaction time per pulse increases. Fig 2 shows the evolution of short-pulse laser interaction zone after 20 ns of laser irradiation with fluence 3 mJ/cm2 (laser intensity ~ 1013 W/m2), as compared to the previous results, the melt layer is extremely thin (< 2 \u00b5m). The main mechanism of material removal is vaporization, which is depicted by escaping vapor velocity arrow plot (max velocity \u0334100 m/s). The HAZ and melt zones are considerably small as compared to (\u03bcs-ms) drilling. From Fig. 1 and 2, it can be inferred that pulse interaction time and laser intensity largely affects the role of melt pool dynamics in laser drilling. Thus precise understanding of hole formation during laser drilling with appropriate spatial and temporal resolution was possible with the developed model. Further, different regimes such as material removal by vaporization, melt expulsion, melt shadowing and hole blockage have been elucidated. The experimental observations and its comparison with simulation data can be found in [11]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001020_s00170-015-7033-2-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001020_s00170-015-7033-2-Figure7-1.png", "caption": "Fig. 7 Grinding of helix groove and rake face", "texts": [ " Then the workpiece moves along z-axis to determine the grinding azimuth. Meanwhile, the grinding wheel feeds along the positive direction of z-axis and rotates around y-axis to meet the helix angle. 2. The workpiece moves to the required distance along yaxis. A point i on the grinding face rotates around yp-axis by \u03b8, and the grinding wheel rotates around yp-axis by \u03b4, as shown in Fig. 6. 3. The workpiece is feed along the direction of X direction tangent to the cutting edge when rotating around a-axis to form the rake face, as shown in Fig. 7.Fig. 14 Accuracy test of cutting edge After the grinding, the grinding wheel moves up along yaxis and moves back the required distance along x-axis. The workpiece moves a distance equals to the cutter\u2019s groove spacing and rotates around a-axis. Then the grinding wheel feeds along X direction tangent to the cutting edge to complete grinding another cutter\u2019s groove. The grinding wheel\u2019s position when grinding flank of the cutting edge is determined by the flank angle as well as the unit tangent vector and binormal vector" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003684_j.matpr.2021.05.466-Figure14-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003684_j.matpr.2021.05.466-Figure14-1.png", "caption": "Fig. 14. Fixed support.", "texts": [ " Initially, low carbon steel was assigned. Then, meshing was created. Fig. 12 shows the example of meshing done in a hemispherical end cylinder. The size of the meshing was kept at 5 mm. The shape of the mesh was tetrahedral. The nodes and elements of the various 3D models are described in Table 1. After Meshing, boundary conditions were applied. Fixed support was applied at the bottom face of the saddle as shown in Fig. 13. For Longitudinal Stress, 8 bar pressure was applied at the ends or heads of the cylinder (Fig. 14). For Circumferential Stress, 8 bar pressure was applied to the side faces or middle portion of the cylinders (Fig. 15). Although, the results of longitudinal stress were evaluated because the longitudinal stress alone act on the ends of cylinders. This project aims to study the effect of ends with constant thickness. Figs. 16, 17, 18, and 19 shows the result of longitudinal stressinduced in Elliptical, Hemispherical, Torispherical, and Plain formed head. Fig. 23. Circumferential stress for low carbon steel" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000588_9783527822201.ch15-Figure15.9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000588_9783527822201.ch15-Figure15.9-1.png", "caption": "Figure 15.9 Overview of the bending parameters for method B: (a) side view with typical curvature achieved by a deposition of a different amount of layers (n1, n2) with equal thickness (h); (b) top view of the plotting area depicting relevant angles and sectors for the deposition process; and (c) typical resulting structure.", "texts": [ " To address these points, after finishing the deposition in the previous direction, it is necessary to keep the plotting mechanism rotating for a few degrees without feeding. And then to stop for a few seconds also the plotting to guarantee the cooling and fixing of the previous layer, before restarting deposition in the reverse direction. A combination of complete layers with half deposited layers generates an asymmetry that results in bending of the structure. The relevant parameters for the growing task are (Figure 15.9) the bending direction (\ud835\udf03), which is the central angle of the sector without material; the amplitude of the sector with material named the differential deposition sector (\u0394D) that has the center in the direction opposite to the bending angle (\ud835\udf03 + 180\u2218); the angle before the inversion process between the stop of the feeding and the stop of the plotting, named the extra angle (EA); and the deposition ratio (K ), the number of full deposition circle between a single differential deposition. The differential deposition task can be summarized in three steps: (1) Plotting and feeding motor are turned on (Figure 15" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.4-1.png", "caption": "FIGURE 2.4 Forces and moments, acting on tire, speeds, and slip speeds.", "texts": [ " However, for an engineer to use practical tools to describe steady-state tire behavior, the underlying physical phenomena must first be understood. For that reason, physical tire models are discussed in Section 2.7. We will treat two types of models, the brush model (which describes the local deflection in the contact area by linear springs) and the brush-string model where the belt deformation is also accounted for. These models will allow us to examine the local contact phenomena between tire and road. A tire is schematically shown in Figure 2.4, with all the output quantities (forces and moment) and speeds indicated. Note that the z-axis is chosen in the downward direction. There are three forces and three moments acting on the tire (the output quantities): Forces Fx : Brake/drive force Fy : Lateral (cornering) force Fz : Tire load (to carry the vehicle weight) Moments Mx : Overturning moment My : Moment about the wheel axis (drive/brake torque) Mz : Self-aligning moment These forces and moments depend on a number of input quantities, which will be discussed in the subsequent sections: A tire travels with a horizontal velocity V with components Vx and Vy in longitudinal and lateral direction, respectively. Due to brake or drive torque and cornering forces, slip will occur, which means that the tire slides with nonzero speed over the surface. The corresponding slip speeds Vsx and Vsy are shown in Figure 2.4 as well. Note that the slip quantities tan(\u03b1) and \u03ba, introduced previously, correspond to the negative ratios of slip speed and forward speed in x direction. The tire rolls over the surface with an angular speed \u03a9, leading to the rolling speed: Vr 5\u03a9 Re \u00f02:1\u00de where Re is the effective rolling radius of the free rolling tire. For a free rolling wheel (zero slip speed), the rolling speed coincides with Vx; therefore, the effective rolling radius is defined as the ratio between Vx and \u03a9 under these conditions", " Note that the resulting vertical contact force acts slightly in front of the wheel center, meaning that the pneumatic trail may become negative for excessive sliding. Hence, we obtain a side force Fy(\u03b1), starting at small values, at \u03b15 0, and growing to a maximum value (with the full contact area in sliding conditions, and therefore, Fy5\u03bc Fz), whereas the pneumatic trail tp(\u03b1) starts at large values and reduces to small values with even negative values for excessive slip. Pneumatic trail times side force yields the self-aligning torque (or moment) Mz, introduced in Section 2.1 (see Figure 2.4). This torque is called self-aligning because it tries to orient the tire in the speed direction. It works against the lateral deformation due to the lateral force. For pure lateral slip (no braking or driving), the self-aligning torque can be described as follows Mz\u00f0\u03b1\u00de52tp\u00f0\u03b1\u00de Fy\u00f0\u03b1\u00de1Mzr\u00f0\u03b1\u00de \u00f02:41\u00de for residual torque Mzr, a small torque that results from inaccuracies in the tire design that rapidly decreases in absolute value with increasing slip angle. When the tire experiences a brake or drive force in combination with lateral slip (we call that a situation of combined slip), the lateral deflection from the symmetry plane times the longitudinal force will contribute to the aligning torque, as will be discussed further in the next sections" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000790_rcar.2018.8621743-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000790_rcar.2018.8621743-Figure3-1.png", "caption": "Fig. 3. The coordinate of robot Fig. 4. The global coordinate", "texts": [ " 1 shows the mecanum wheel which includes the hub and the passive rollers. From the Fig. 2, the following equation can be calculated[ v\u2032ix v\u2032iy ] = [ 0 sin\u03b1i r cos\u03b1i ][ \u03c9i vir ] = Ki1 [ \u03c9i vir ] (7) where all parameters are in the wheel coordinate, v\u2032ix, v\u2032iy, \u03c9i are the wheel speed and angular velocity. vir represents the norm velocity of the wheel. r represents the radius of the wheel. It is necessary to translate v\u2032ix, v\u2032iy in the x\u2032oy\u2032 coordinate to vix, viy in the xoy coordinate. From the Fig. 3, the equation in the robot coordinate as follows:[ vix viy ] = [ cos\u03b8i \u2212sin\u03b8i sin\u03b8i cos\u03b8i ][ v\u2032ix v\u2032iy ] = Ki2Ki1 [ \u03c9i vir ] lix = licos\u03b2iliy = lisin\u03b2i (8) where Ki2 is the transition matrix. li is the distance of the two origin points between the wheel coordinate and the robot coordinate. According to geometric properties[ vix viy ] = [ 1 0 \u2212liy 0 1 lix ][ vx vy \u03c9 ] = Ki3 [ vx vy \u03c9 ] (9) The transformation relationship of robot speed in different coordinate systems is obtained by Eqs.(9). Combining Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003440_j.mechmachtheory.2021.104297-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003440_j.mechmachtheory.2021.104297-Figure1-1.png", "caption": "Fig. 1. The 3\u20133 Gough-Stewart platform with the elastically supported moving platform.", "texts": [ ", both Schur complements are undefined, then the robot singular configurations can be formulated only as det[ K ] = 0. The results are summarized in Table 1 . To extend the results relating to collinear stiffness values obtained in works [26-31] , this paper is focused on the singularity problem formulation in terms of the Schur complements. As methodological examples, the Hunt-type singular configuration of the Gough-Stewart platform (GSP) [34] and the RRR plane manipulator are considered. In Fig. 1 , one of the most popular architecture (3\u20133 version) of the GSP is shown. The kinetostatics of the GSP mechanism is modeled as the rigid moving platform connected with the rigid base through six elastic limbs. In this model, the limb is considered as a chain composing three linear elastic elements: an actuator of the translational motion, and two spherical joints connecting the limb with the moving platform and the base. Since the GSP presents a parallel-kinematics system, its stiffness matrix is expressed through equation K = J T K limb J (67) where K limb is the diagonal matrix combining stiffness values k i of the limbs ( i = 1, 2, \u2026, 6), and J is the Jacobian matrix, K lumb = Diag [ k 1 , k 2 , " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002427_b978-0-12-420017-3.00003-7-Figure4.64-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002427_b978-0-12-420017-3.00003-7-Figure4.64-1.png", "caption": "Figure 4.64 General mechanism of WSP-based probes\u2019 reaction to H2S showing release of the fluorophore.", "texts": [ " The turn-off mechanism exploits the Michael addition mediated reductive cyclization of aldehyde rhodamine conjugate as shown in Fig. 4.63. The availability of SH after reaction to H2S leads to intramolecular cyclization, which makes this a unique strategy to distinguish between H2S and other primary thiols (RSHs). High sensitivity to H2S compared to Cys and GSH was observed with WSP-based probes, which use dual-electrophilic centers\u2014i.e., a disulfide that reacts with the H2S and an ester linker group that can release the fluorophore on cyclization, according to Fig. 4.64.204 By linking 2,4-dinitrophenyl (DNP) to a fluorophore via an ether linkage as DNPPCy (Fig. 4.65) allows for a different strategy in H2S detection in which the nucleophilic character of the thiol is exploited for thiolysis. Nucleophilic aromatic substitution reaction with H2S leads to the liberation of the fluorophore as a phenolate that has a NIR emission property that allows for ratiometric measurement of H2S. However, DNPPCy exhibits a slight fluorescence response to GSH and Cys.205 Instead of an ether linker group, SBODIPY-DNP (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003249_tte.2021.3049466-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003249_tte.2021.3049466-Figure6-1.png", "caption": "Fig. 6. Cross section of the rotation process for the deflection teeth.", "texts": [ " However, the hybrid slots will create an additional problem, namely unequal slots, as shown in Fig. 4. It possesses twotype slots with different areas (slots A and B) for each stator, and the area of slot A is always larger than that of slot B if all the slot-pitch angles are the same. It will reduce the utilization rate of slots A because these two-type slots share the same windings. In order to improve the utilization rate of slots, the deflection teeth, which rotate partial teeth toward adjacent phases by a certain angle, are proposed. Fig. 6 shows the cross section of the rotation process for deflection teeth. The teeth with dotted lines are those that are designed initially with an equal slot-pitch angle before rotation, and the teeth with the solid line are those that are designed for unequal slot-pitch angles after rotation. \u03b8i and \u03b8o are corresponding rotation angles of two stators. Fig. 7 shows the performances of HS-RDPM versus \u03b8i and \u03b8o. It can be seen that the torque increases first with \u03b8i or \u03b8o before decreasing, and reaches the top when \u03b8i and \u03b8o are 0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002762_j.euromechsol.2020.104125-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002762_j.euromechsol.2020.104125-Figure6-1.png", "caption": "Fig. 6. (a) Test bearing pedestal (b) Angular markings on pedestal (c) acceleration in x and y direction of measurement (d) acceleration obtained after vector addition.", "texts": [ " By employing this analogy, the variation of load in load zone and its consideration in defect size estimation is taken care of, and it is applicable for both outer race defect as well as inner race defect. With this analogy, the force corresponding to various peaks shown in Fig. 5(b) are enlisted in Table 1. To verify the process of obtaining force from acceleration peak magnitude as described above, experiments were conducted for outer race defect by keeping the defect at different locations in the load zone. For this, angular markings were marked on the lower half of the pedestal, as shown in Fig. 6(a) and (b). The outer race was so oriented that the defect was placed at a different angular location in the load zone for each experiment. For each defect orientation, experiments were conducted at shaft speeds from 300 rpm to 1200 rpm with increments of 150 rpm. The acceleration was recorded in horizontal (x-direction) and vertical direction (y-direction) by two sensors, as shown in Fig. 6(a). Then, the vector addition of the acceleration in horizontal and vertical was done to get the effective acceleration. The acceleration in x and y direction for ORD1 (refer Table 2 in section 3) at 1200 rpm and defect positioned at 6 o\u2019clock are shown in Fig. 6(c) and the resultant acceleration is shown in Fig. 6(d) for 0.1 s data. From the resultant acceleration, all the peak accelerations are obtained by the process described in section 2.3.1, and the mean of all peak accelerations is obtained. Due to dynamic conditions, for a given defect orientation, load, and speed, the peak acceleration at the rolling element-defect interaction is different at different time instances (visible in Fig. 6(d)). Hence, to eliminate this variation, the mean of all the peaks of resultant acceleration is used for the analysis of force-peak acceleration correlation. The mean peak acceleration for ORD1 at 1200 rpm and different angular location of defect in load zone is shown in Fig. 7(a). The levels shown by horizontal dotted lines indicate different thresholds of mean peak acceleration, topmost being maximum acceleration amax, and lower most being the 0.2amax. As mentioned earlier, the acceleration value lying in a particular threshold band is rounded to the higher threshold limit value", " Since cross-correlation indicates the similarity of two signals as a function of the displacement of one relative to the other, the maximum value of cross-correlation given by Eq. (27) is obtained when the starting point of the target signal coincides with point B. Hence, the time instance having the highest magnitude of cross-correlation is chosen as point B, and the time corresponding to B is termed as TB. The cross-correlation of the target signal and SOI is shown in Fig. 9(d) for demonstration. In the studies (Jena and Panigrahi, 2014; Khanam et al., 2014; Sawalhi and Randall, 2011; Singh and Kumar, 2013), the point corresponding to Tp (Fig. 6 (a)) was considered the exit point (point C). To obtain the point C, the method employed in Ref (Chen and Kurfess, 2019). was to locate the transition point on the envelope of the signal obtained after a second derivate of the acceleration signal. Pre-whitening combined with multi-resolution analysis and Hilbert transform was used in Ref (Sawalhi and Randall, 2011). to obtain point C, while Wavelet decomposition was used in Ref (Jena and Panigrahi, 2014; Khanam et al., 2014; Singh and Kumar, 2013)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000588_9783527822201.ch15-Figure15.6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000588_9783527822201.ch15-Figure15.6-1.png", "caption": "Figure 15.6 (a) Closed view of the growing robot, with main components and (b) inside view of the deposition head reproducing the 3D printer-like mechanism.", "texts": [ " Such robotic mechanism allows the implementation and testing of plant-inspired behaviors. Taking inspiration from the ability of plants to adapt their shape and structure while growing, in [14] an onboard 3D printing mechanism for robotic roots has been investigated. The system is designed by exploiting the classic approach of FDM and integrates a customized 3D printer inside the artificial root to extrude thermoplastic material on purpose. The growing system is composed by a tubular body, a growing head, and a sensorized tip that controls the robot behaviors (Figure 15.6a). The principle of the growing head is similar to the one described in Section 15.4.4, but in this case the 3D printer-like system is composed of an extruder and a plotting unit (Figure 15.6b). The extruder unit includes a feeder mechanism that pulls a filament from outside, and then pushes it through a guiding tube toward a heater, which in turn externally fuses the filament that can be extruded from a nozzle. The plotting unit provides a rotational motion to the growing head generating a circular deposition of the extruded material. The growing head is then interfaced with the root tubular body through four flexible clamping fingers installed on its circumferential sides to prevent the rotation of the growing mechanism inside the printed tubular body" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure2-1.png", "caption": "Fig. 2. To the differential spatial displacement of body n in Eq. (2).", "texts": [ " , n, (1) with the unit vectors ai, |ai| = 1, of the screw axes, the moments r\u0303i ai of the vectors ai with respect to a common reference point, here point On on body n, and the pitches hi of the helical joints. In the actual position of the kinematical chain, a differential relative rotation dqi of the ith helical joint generates a differential spatial displacement of body n relative to body 0 comprising the differential rotation vector dxn(i) \u2261 aidqi and the differential translation vector drn(i) of point On, see Fig. 2, [ dxn(i) drn(i) ] = [ ai r\u0303iai + hiai ] \ufe38 \ufe37\ufe37 \ufe38 a\u0302i dqi \u2261 [ ai aei ] dqi. (2) The resulting differential spatial displacement of body n due to the independent differential screw motions a\u0302i dqi, i = 1, . . . , n, is then obtained by superposition, [ dxn drn ] = [ a1 a2(q) . . . an(q) ae1 ae2(q) . . . aen(q) ] \u23a1\u23a2\u23a2\u23a2\u23a2\u23a2\u23a2\u23a3 dq1 dq2 ... dqn\u22121 dqn \u23a4\u23a5\u23a5\u23a5\u23a5\u23a5\u23a5\u23a6 \u2192 dx\u0302n = G(q) dq. (3) As the axis of the first helical joint of the nH chain is in the base body 0, its screw coordinates of a1, ae1 are constant, while the screw coordinates of the other joints depend on the joint coordinates q", " a\u0302n\u22121,0 ] , A\u2032 0 = [ da\u03021 ds \u2223\u2223\u2223\u2223 s0 da\u03022(s) ds \u2223\u2223\u2223\u2223 s0 . . . da\u0302n\u22121(s) ds \u2223\u2223\u2223\u2223 s0 ] \u2261 [ 0 a\u0302\u2032 20 . . . a\u0302\u2032 n\u22121,0 ] , A\u2032\u2032 0 = \u23a1\u23a3 d2a\u0302\u2032\u2032 1 ds2 \u2223\u2223\u2223\u2223\u2223 d2a\u0302\u2032\u2032 2(s) ds2 \u2223\u2223\u2223\u2223\u2223 s0 . . . d2a\u0302\u2032\u2032 n\u22121(s) ds2 \u2223\u2223\u2223\u2223\u2223 s0 \u23a4\u23a6 \u2261 [ 0 a\u0302\u2032\u2032 20 . . . a\u0302\u2032\u2032 n\u22121,0 ] , ... and with the derivatives of k(s) k0 = k(s0), k\u2032 0 = dk(s) ds \u2223\u2223\u2223\u2223 s0 , k\u2032\u2032 0 = d2k(s) ds2 \u2223\u2223\u2223\u2223\u2223 s0 , . . . To formulate the derivative a\u0302\u2032 k of a helical joint axis a\u0302k with respect to s, the differential of a\u0302k due to increments of the joint coordinates dqi in the open nH chain according to Fig. 2 is expressed as the sum of the dual vector products, see Eq. (77) in the Appendix, da\u0302k = k\u22121\u2211 i=1 \u02dc\u0302ai a\u0302k dqi. (15) For the nH mechanism the derivative of a\u0302k with respect to s becomes da\u0302k(s) ds = k\u22121\u2211 i=1 \u02dc\u0302ai a\u0302k dqi ds (16) and with Eq. (13) a\u0302\u2032 k = \u2212 k\u22121\u2211 i=1 \u02dc\u0302ai a\u0302k ki . (17) The derivatives of the n joint screws a\u0302i of the nH mechanism with respect to s then are (the screw axes a\u03021, a\u0302n are fixed) a\u0302\u2032 1 = 0 , a\u0302\u2032 2 = \u2212\u02dc\u0302a1 a\u03022 k1 , ... a\u0302\u2032 n\u22121 = \u2212\u02dc\u0302a1 a\u0302n\u22121 k1 \u2212 \u02dc\u0302a2 a\u0302n\u22121 k2 \u2212 . . . \u2212 \u02dc\u0302an\u22122 a\u0302n\u22121 kn\u22122 , a\u0302\u2032 n = 0 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001137_ipec.2014.6869708-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001137_ipec.2014.6869708-Figure2-1.png", "caption": "Fig. 2. Structure of 12/14 hybrid stator pole type BLSRM.", "texts": [ " Similarly, the suspending forces for the y-direction are generated by the currents iyp and iyn, which flow in suspending force poles Pyp and Pym respectively. Meanwhile, to get a continuous suspending force, the suspending force pole arc needs to be wider than one rotor pole pitch. In that way, the aligned area between the suspending force and the rotor pole is always the same which may decrease the effect of suspending force current to the torque. B. 12114 Hybrid Stator Pole Type BLSRM A novel 12/14 hybrid stator pole type BLSRM is proposed in Fig. 2. The proposed structure is similar to the structure of the 8/1 0 type. There are also torque and suspending force poles in the 12114 type. But in the 12114 type short flux paths are taken and no flux reversal exists in the stator core. Windings on the torque poles PA/, PA2, P A3 and PM are connected in series to construct phase A, and windings on the torque poles PRJ, Pm, PRJ and PR\ufffd are connected in series to construct phase B. Similar to the 8110 type, the windings on the suspending force poles Pxp' Pxm Pyp and Pyn are independently controlled to construct four suspending forces in the x- and y directions" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003249_tte.2021.3049466-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003249_tte.2021.3049466-Figure8-1.png", "caption": "Fig. 8. Cross sections. (a) CS-FTPM. (b) SC-RDPM. (c) DC-RDPM.", "texts": [ "2 mechanical degrees are selected for \u03b8i and \u03b8o, respectively, in this part, and the M/L is 2.63%, while k is 1.144. AND INTERIM STRUCTURES According to the results of Section II, there are three optimizations in HS-RDPM, including dual stators, optimization of torque ripple, and optimization of phase isolation. In order to prove the effectiveness of these optimizations, another three 12-slot/ten-pole topologies, namely conventional surface FTPM (CS-FTPM), single-layer concentrated-winding RDPM (SC-RDPM), and double-layer concentrated-winding RDPM (DC-RDPM), are designed in Fig. 8. The common parameters are listed in Table I. Fig. 8(a) shows the cross section of CS-FTPM. As stated in Section I, the SFCW and surface PMs are used to achieve the physical, thermal, and magnetic isolations. The DNNs are employed to increase the self-inductance for limiting the short-circuit current. The cross section of SC-RDPM is shown in Fig. 8(b). It has combined the characteristics of CS-FTPM Authorized licensed use limited to: Dedan Kimathi University of Technology. Downloaded on May 18,2021 at 11:03:36 UTC from IEEE Xplore. Restrictions apply. with dual stators directly. It means that the SFCW, surface PMs, and DNNs, are applied to SC-RDPM. To fit in with the feature of SC-RDPM, the PMs are divided into two layers. Fig. 8(c) shows the cross section of DC-RDPM, in which the optimizations of torque ripples, including DFCW, optimized rotor, and staggered stators, are adopted. Thus, the above three topologies own no or partial optimizations of HS-RDPM, as summarized in Table II: 1) CS-FTPM is based on conventional FTPM, without dual stators, optimizations of torque ripple, and phase isolation; 2) SC-RDPM is based on conventional FTPM and dual stators, without optimizations of torque ripple and phase isolation; 3) DC-RDPM is based on dual stators and optimization of torque ripple, without optimization of phase isolation; and 4) HS-RDPM possesses all of above optimizations" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002154_imcec.2016.7867416-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002154_imcec.2016.7867416-Figure2-1.png", "caption": "Fig. 2. The key geometric parameters ofUAV", "texts": [ " The GPS coordinates of GPS module acquisition UAV information. The image acquisition module is to collect the image in the process of UAV flight. The sensor module acquires status parameters ofUAV during the process of flight. (II') ( Current sensor ) .... .... ( ZigBee module ) U ( Wind sensor ) .... A .... ( GPS module ) V ( Power ) .... ....( Camera ) Fig. l. The structure diagram ofUAV platform III. Experimental materials and methods This research adopts the single screw after push type fixed-wing carrier as shown in Fig. 2, the fuselage materials for foamed polypropylene (EPP), its wingspan is 172 cm, and the captain is 112 cm, 40 cm tail plane. Communication module parameters: CC2530 wireless module of TI Company, the carrier frequency is 2.4 GHz, adopt 3 dBi omni-directional sucker antenna; Heaven and earth fly remote control 8 channels. Power system using the new west company 2217 - KVIIOO 40A brush less motor, brush less electronic governor, the 12 V power dc power supply, the length of254 mm quality for 7g double propellers" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003440_j.mechmachtheory.2021.104297-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003440_j.mechmachtheory.2021.104297-Figure9-1.png", "caption": "Fig. 9. Kinematics of the RRR planar manipulator.", "texts": [ " 7 c, these determinants are made with determinant of the Jacobian matrix shown in Fig. 4 . Note that the values of the three determinants are numerically reduced to focus on the fact that all their zeros are identical. Hence, the determinants of the 3 \u00d7 3 Schur complements can be successfully used for fixation of singular configuration instead of the 6 \u00d7 6 Jacobian and stiffness/compliance matrices. In Fig. 8 , the Schur complements eigenvalues are shown. Obviously, there exist the zero values of the eigenvalues at the singular-configuration-associated points. In Fig. 9 , the RRR serial planar manipulator is shown. The 3 \u00d7 3 Jacobian matrix is defined as follows: J = [ j 11 j 12 j 13 j 21 j 22 j 23 1 1 1 ] (76) j 11 = - a 1 sin \u03d51 - a 2 sin ( \u03d51 + \u03d52 ) - a 3 sin ( \u03d51 + \u03d52 + \u03d53 ) j 12 = - a 2 sin ( \u03d51 + \u03d52 ) - a 3 sin ( \u03d51 + \u03d52 + \u03d53 ) j 13 = - a 3 sin ( \u03d51 + \u03d52 + \u03d53 ) j 21 = a 1 cos \u03d51 + a 2 cos ( \u03d51 + \u03d52 ) + a 3 cos ( \u03d51 + \u03d52 + \u03d53 ) j 22 = a 2 cos ( \u03d51 + \u03d52 ) + a 3 cos ( \u03d51 + \u03d52 + \u03d53 ) j 23 = a 3 cos ( \u03d51 + \u03d52 + \u03d53 ) where a 1 , a 2 , and a 3 are the link lengths; \u03d51 , \u03d52 , and \u03d53 are the angles of the link rotation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003598_tmag.2021.3078841-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003598_tmag.2021.3078841-Figure10-1.png", "caption": "Fig. 10. The magnetic flux density vector distribution in the unsaturated 3-D", "texts": [ " It can be seen that the results of the analytical model have good agreement with those of the unsaturated 3-D FEA model, validating the proposed analytical model. Furthermore, the calculation error is dramatically reduced compared with Fig. 6 (a), which means that ignoring saturation and stator slot can lead to significant calculation error. It should be also noted that there remains obvious calculation error between the unsaturated 3-D FEA and the analytical model in Tm. The magnetic flux density vector distribution in the unsaturated 3-D FEA model is illustrated in Fig. 10. Although saturation and slot effect are both eliminated in the unsaturated 3-D FEA model, it can be clearly observed that severe leakage flux occurs in the adjacent air space near the inner and outer radii due to end effect. The proposed analytical model cannot take such end effect into account, resulting in inevitable calculation error. FEA model. Although there remains inevitable calculation error with the analytical model, the analytical and 3-D FEA results follow the same changing trends and give the same optimal values of the ferromagnetic pole-pieces pole-arc coefficient and the PMs thickness" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001774_j.ijleo.2016.04.037-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001774_j.ijleo.2016.04.037-Figure1-1.png", "caption": "Fig. 1. Typical planar engagement geometry.", "texts": [ " The rest of this paper is organized as follows: Section 2 is dedicated to formulating the missile-target engagement inematics. The SFS-SOSM control algorithm is derived and its FTMSP convergence is proved in Section 3. A smooth guidance aw based on SFS-SOSM control is presented in Section 4. In Section 5, the performances of designed guidance law are verified ia several computer simulations. . Planar engagement model and intercept strategy .1. Problem formulation Consider the planar homing case that the missile moves within the horizontal plane, a typical engagement scenario is resented in Fig. 1. The planar missile-target engagement kinematics can be easily derived as r\u0307 = VT cos(q \u2212 T ) \u2212 VM cos(q \u2212 M) (1) rq\u0307 = \u2212VT sin(q \u2212 T ) + VM sin(q \u2212 M) (2) here q is the LOS angle; r is the range along LOS; VT and VM are target and missile velocity; T and M are target aspect ngle and missile heading angle, respectively; T = q \u2212 T and M = q \u2212 M are target and missile\u2019s lead angle. Denote Vr = r\u0307, Vq = rq\u0307 and substitute them into Eqs. (1) and (2), then differentiating both sides of Eqs. (1) and (2) with espect to time t to get V\u0307 r = q\u0307[\u2212VT sin(q \u2212 T ) + VM sin(q \u2212 M)] + [V\u0307T cos(q \u2212 T ) + VT \u0307T sin(q \u2212 T )] \u2212 [V\u0307M cos(q \u2212 M) + VM\u0307M sin(q \u2212 M)] (3) V\u0307 q = \u2212q\u0307[VT cos(q \u2212 T ) \u2212 VM cos(q \u2212 M)] + [VT \u0307T cos(q \u2212 T ) \u2212 V\u0307T sin(q \u2212 T )] \u2212 [VM\u0307M cos(q \u2212 M) \u2212 V\u0307M sin(q \u2212 M)] (4) Let aT,r = V\u0307T cos(q \u2212 T ) + aT sin(q \u2212 T ) (5) aM,r = V\u0307M cos(q \u2212 M) + aM sin(q \u2212 M) (6) aT,q = aT cos(q \u2212 T ) \u2212 V\u0307T sin(q \u2212 T ) (7) aM,q = aM cos(q \u2212 M) \u2212 V\u0307M sin(q \u2212 M) (8) where aM = VM\u0307M , aT = VT \u0307T are normal acceleration of missile and target, respectively; aT,r , aT,q, aM,r , aM,q are target and missile acceleration along and orthogonal to LOS, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003519_s10846-021-01359-5-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003519_s10846-021-01359-5-Figure1-1.png", "caption": "Fig. 1 The remotely controlled leader robot(left) guides the fully autonomous follower robot(right) with a payload clamped on both of them. The follower senses the force through the payload and computes an obstacle-avoidance trajectory", "texts": [ " Furthermore, a key aspect of making a mobile robot autonomous is by adding an obstacle-avoidance capability to it. Hence, the follower robot computes a coordinated obstacle-avoidance motion based on the leader force. The payload clamped on these robots is used to transmit the leader force to the follower robot. An additional degree of freedom is provided as a constrained relative motion between the payload and the follower robot towards the applied force. This force is sensed as a direct co-relation to displacement produced on this DoF. A visualization of this set-up can be seen in Fig. 1. The rest of the paper is organized as follows: a detailed background and related work are discussed, followed by the adopted methodological approach. In this section, the basic assumptions of the study are presented. Additionally, the system is formulated into four specific cases, which are further elaborated. The proposed coordination architecture of the collaborative manipulation and the follower robot\u2019s control algorithm are described under it. Furthermore, Kinematic formulation of the MeWBot is stated and the quantified overall system is represented mathematically" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002391_s41315-020-00127-2-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002391_s41315-020-00127-2-Figure2-1.png", "caption": "Fig. 2 Visual SLAM for endoscope navigation and mapping", "texts": [ " By combining the pre-operative CT scan data with registration, the scale of both the endoscope trajectory and 3D reconstructed map can be recovered and the operative guidance can be implemented. As far as we know, our method is well-designed for complex oral scenes and it is also the first one to solve the transoral endoscope navigation problem with Visual SLAM technique, making full use of intra-operative and pre-operative information in the meanwhile. 3.1 Transoral SLAM to\u00a0perform navigation and\u00a0mapping In our transoral endoscope navigation system, we adopt ORBSLAM to realize endoscope localization and mapping, as shown in Fig.\u00a02, which performs better than other wellknown SLAM techniques, such as PTAM (Klein and Murray 2007) and LSD-SLAM (Engel et\u00a0al. 2014). Many state-ofthe-art techniques are integrated into this framework and the ORB descriptor makes it a real-time system. In addition, in MIS ORBSLAM can still keep robust tracking and mapping against human in-vivo tissues where slight deformation and partial occlusions may take place. As we all know, the oral cavity is a complicated part of the human body, which is consisted of several components: tongue, hard palate, lips, gingivae and teeth" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001816_j.proeng.2016.05.097-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001816_j.proeng.2016.05.097-Figure5-1.png", "caption": "Fig. 5. Time response and orbit plots of a stable and unstable system", "texts": [ " The threshold speed for rotor supported on powder lubricated bearings is more than three times of that when the bearings are oil lubricated. Also, the threshold speed is more for rigid rotor than that of a flexible rotor as observed by Rao [18]. The response obtained by moving the journal center to the bearing center and then releasing it reflect the stability of the system known as position perturbation [19]. For position perturbation, the time response and the trajectory of the journal center for a flexible shaft supported on powder lubricated bearings are shown in Fig. 5. Two cases are shown; (a) a stable system at 20000 rpm, and (b) an unstable system at 27000 rpm. For a stable system the amplitude of displacement decreases with time and the journal goes back to its steady state position. For an unstable system the amplitude of displacement increases with time and the journal goes far and far away from the equilibrium position. A powder lubricated journal bearing is investigated for the evaluation of its dynamic characteristics. Stiffness and damping coefficients are obtained and stability limits of spin speed of a rigid as well as flexible rotor consisting of a central disc are investigated and compared with the case when the bearings are oil lubricated" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002931_tmech.2020.3036765-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002931_tmech.2020.3036765-Figure1-1.png", "caption": "Fig. 1. CMG description.", "texts": [ " Section III details the proposed nonlinear controller design for the CMG MIMO tracking task problem and the stability of the overall closed-loop system is analyzed by invoking both Lyapunov stability and LaSalle\u2019s invariance principle. Next, the proposed nonlinear controller is evaluated in Section IV, in which numerical simulations and real-time practical experiments are performed for two considered tests and its results are discussed, along with a performance analysis of its experiment results by overlapping it with the ones obtained from the nonlinear controllers of [22] and [23]. Finally, Section V provides concluding remarks. The ECP CMG unit, described in Fig. 1, is an underactuated electromechanical plant with two actuators (DC motors #1 and #2) and four degrees-of-freedom (DoF) represented by the angular positions (\u03b8i(t)) of bodies A, B, C, and D, that rotate around axes i = 1, 2, 3, 4, respectively, being measured by incremental encoders A, B, C, and D. Motors #1 and #2 apply torques T1(t) and T2(t) at axes #4 and #3, respectively, rotating bodies D and C. On the other hand, bodies B and A are rotated only by intrinsic reaction torques. The rotation on Body B is mainly due to reaction of torque T1, obeying the principle of angular momentum conservation, and body A is rotated by the gyroscopic torque generated when the torque T2 is applied while body D is rotating", " Remark 1: Body D, called rotor, is an axisymmetric brass disk responsible for generating the angular momentum that is essential for the CMG operation, thus its velocity must not be zero. Remark 2: Note that this system is underactuated and has two subsystems: an unactuated one, composed of bodies A and B dynamics, and an actuated one, composed of bodies C and D dynamics. The main goal of the controller design herein is to perform the tracking task of the unactuated states (\u03b81 and \u03b82) with the system control action composed of both motor torques (T1 and T2). The CMG zero position orientation (\u03b8i = 0) is depicted in Fig. 1, where body C is oriented perpendicular to body B, which, in turn, is oriented perpendicular to body A. This is important because the CMG operates respecting the following assumptions: Assumption 1: Due to the existence of a singularity at \u03b82 = \u00b1\u03c0/2 rad (90\u25e6), which is known as a gimbal lock [30], the body B angular position operates within the interval of approximately \u00b1\u03c0/3 rad (60\u25e6), i.e., \u2212 \u03c0/3 \u2264 \u03b82(t) \u2264 \u03c0/3, \u2200t \u2265 0. (1) Assumption 2: The body C angular position is always within the interval of \u00b1\u03c0/2 rad (90\u25e6) during its operation, due to physical constraints, i" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001392_elektro.2014.6848921-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001392_elektro.2014.6848921-Figure2-1.png", "caption": "Fig. 2. Schematic representation of dynamic eccentricity", "texts": [ " Eccentricity in Electrical Machines Two different types of eccentricity can be distinguished: \u2022 static eccentricity (SE) \u2022 dynamic eccentricity ( DE) Both of them cause unbalanced magnetic pull (UMP) , i.e. the radial force acting in direction of minimal air-gap. While in case of SE the position of minimal air-gap is constant in time 978-1-4799-3721-9/14/$31.00 mOl41EEE 375 and the value and direction of UMP can be considered invariable in the first approximation, in case of DE the position of minimal air-gap and direction of UMP varies in time. This fact is obvious in Fig. I and Fig. 2. SE is usually characterized by perfectly rigid shaft which is set into the stator with slight misalignment. This state can cause the shaft deflection, particularly in case of long shaft. However this deflection has constant position as stated before. On the other hand, DE is mostly caused by the shaft deflection given during the production or by thermal imbalance. Consequently the shaft oscillation can occur. Besides the shaft deflection or oscillation and the stator core or machine's frame vibration, eccentricity can cause other adverse effects" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003411_lra.2021.3062296-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003411_lra.2021.3062296-Figure10-1.png", "caption": "Fig. 10. On the left, the Universal Robot UR-10 manipulator with the endeffector. On the right the, different types of end-effector designs to account for the different dimensions and geometries of the testing objects.", "texts": [ " Therefore, trajectory generation using the SERV OJ command (which allows the user to determine the acceleration profile) were also implemented. These were sent to the manipulator via Real-Time Data Exchange (RTDE) interface in Python, where a set of desired Cartesian poses for the end-effector (representing the desired path) were translated to joint angles to be reached at specified times, depending on the minimum acceleration value as discussed on Section IV. A0 : \u03b80MAX < \u03c0/2 (10) A1 : \u03b80MAX \u2265 \u03c0/2 (11) B1 : \u03b80MAX + \u03b8(tn) \u2265 \u03c0/2 (12) Three different types of end-effector were tested (Fig. 10) with a variety of object shapes and geometries (Fig. 11), many of them clearly not following the assumed characteristics for our approach to help us determine the robustness to our model assumptions. A total of 152 out of 160 trials were successful (95%). Experimental results are shown on Table I. Unsuccessful 2https://www.universal-robots.com/ trials occurred when objects with liquid inside shifted their center-of-mass (Objects #7 and #10), which is not taken into account in this formulation. Additionally, air resistance seemed to have caused Object #9 (which is lightweight) to fall during the accelerated motion" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003301_j.matpr.2020.12.1110-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003301_j.matpr.2020.12.1110-Figure5-1.png", "caption": "Fig. 5. Stress and strain for steel alloy.", "texts": [], "surrounding_texts": [ "The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper." ] }, { "image_filename": "designv11_22_0003411_lra.2021.3062296-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003411_lra.2021.3062296-Figure5-1.png", "caption": "Fig. 5. When losing contact with the ground (without loss of generality considering P2 as frictionless), if the object was to become static, the moment labeling of the system indicates the contacts cannot generate a wrench to balance gravity. However, for an acceleration ad the resulting acceleration aN can balance fc = wg + ac (which comprises gravity and centrifugal acceleration) and be generated by the contacts. The segment G\u2032 w has to lie inside the friction cone at contact P1 (2\u03b5, with \u03b5 = atan\u03bc).", "texts": [ "X\u0308, Y\u0308 , \u03b8\u0307, \u03b8\u0308 are the components of the linear velocity and acceleration (ad = X\u0308i+ Y\u0308 j) and angular velocity and acceleration (\u03b8\u0308 = \u03b8\u0308k) imposed to the object respectively. In general terms, it is intuitive to think of the applied linear and angular accelerationsad, \u03b8\u0308 overcoming the effects of gravity and rotational dynamics (e.g., centrifugal acceleration). We apply a graphical method such as moment labeling to analyze the instant t0 immediately after the object loses contact with the surface (Fig. 5(a)). The left and right edges of the friction cone at contact P1 are F1 l and F1r respectively and contact at P2 is assumed to be frictionless to facilitate visualization of feasible solutions without loss of generality (\u03b5 = 0 at the contact and so F2r = F2 l = F2n becomes the contact normal). The moment labeling of the composite wrench cone of the contacts generate a positive labeled region about which the contacts can only exert positive torque (according to the right hand rule). By definition, no wrench can pass through the region and therefore the scenario in Fig. 5(a) is not stable (since the actuation line of the gravity wrench Lw passes through the positive labeled region) and the object will fall. However, considering Fig. 5(b), if an inertial acceleration ad is imposed to the object, the contacts will be able to apply a vector sum aN = ad \u2212 fc (with fc = wg + ac) to match the wrench of gravity and the applied acceleration. The vector sum can be translated along its line of actuation L\u2032 w without modifying its value, and be generated by a linear combination of the friction cone at P1 (k1 lF1 l + k1rF1r) and normal at P2 (k2F2n). Since this resultant wrench respects the labeled region by generating a positive moment with respect to the positive labeled regions, the system is instantaneously stable", " We can therefore propose candidate acceleration ad and applied torque In\u03b8\u0308 that result in non-negative kij values. 3) Trajectory Generation: The acceleration will change at every time step (according to the rotation angle \u03b8), and an infinite number of trajectories can satisfy the constraints. It is useful to determine the limits of the minimum and maximum ad that satisfy the constraints to determine the boundary trajectories such that any trajectory within these boundaries is a feasible trajectory. The termination condition can also be determined (Fig. 5), since for \u03b8(tn) \u2265 \u03c0/2\u2212 \u03b2 the contact wrenches will be able to span the gravity wrench with no imposed acceleration needed. However, the termination condition at rest requires deceleration as well. From 5, the acceleration ad and torque I\u03b8\u0308 can be chosen to stop rotation and linearly decelerate the object. Although there are infinite choices for the acceleration, by coupling ad and \u03b8 while t \u2264 tn (Fig. 5(b)) it is possible to constrain the system\u2019s trajectory at the termination condition to result in no rotational motion of the object, just translation while decelerating. This is expressed by: ad(t) = X\u0308X\u0302+ Y\u0308 Y\u0302 = h0(cos \u03b8(t)X\u0302+ sin \u03b8(t)Y\u0302), h0 \u2265 0, 0 \u2264 \u03b8(t) \u2264 \u03c0/2 (8) where X\u0302, Y\u0302 are the unit vectors in the world frame axis [X,Y ] respectively. To determine the acceleration magnitudeh0, for every 0 \u2264 \u03b8 \u2264 \u03c0/2, we consider the constraint on the non-negative kij which will determine a lower bound for h0 and therefore to the acceleration ad" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001918_978-3-319-13117-7-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001918_978-3-319-13117-7-Figure8-1.png", "caption": "Fig. 8 . Elastic modulus for the concentration of 22%, the lines represent the lineal regression (LR) for all three temperature values . Their slopes are m12=-0.0013, m16=0.0384 and m20=0.0568.", "texts": [ " closely the \u03a3\u2212\u0394 converter digital filter, which has a sinc3 shape. resulting in a 3dB bandwidth from DC to 512 Hz when for a sampling rate of 2000sps. The CMRR resulted to be of 90 dB with no impedance imbalance. When 51 k\u03a9//47 nF was applied, the CMRR was of only 46 dB. 90 k\u03a9 and 9 k\u03a9 imbalances were tested yielding CMRRs of 56 and 79 dB respectively. A sample of the EMG measurements for setup a is shown in figure 7. An action potential train with increasing fire rate as force increased can be seen. Results for setup b can be seen in figure 8. All hand fingers were extended and flexed, and the resulting EMG signals are shown in the first and second columns respectively. The third column shows the signal resulting from an ulnar flexion of the wrist. The labels of each row correspond to the labels of each electrode shown in Fig. 6. Electrodes were easily put into place, although the ribbon cables used were too stiff and caused movement artifacts. It was observed that a longer time was required for the baseline of the signal to stabilize compared with the case when wetgel electrodes were used, and during that time spurious spikes of high amplitude were produced", " They showed higher low frequency noise and more tendency to produce artifacts than wet-gel electrodes, so the DC-coupled, high dynamic range design was well suited for them. The system parameters were appropriate for EMG measurement. Noise of 1 \u03bcVrms with a gain of 4 provides enough resolution while allowing a DC offset of \u00b1550 mV . Bandwidth is in accordance with preconditioning filters typically used in EMG amplifiers [1]. Because the DRL circuit contributes 30 dB to the common IFMBE Proceedings Vol. 49 Fig. 8: Signals from rigth forearm muscles elicited from: (a) the extension of all hand fingers, (b) finger flexion, (c) wrist ulnar flexion. mode signal attenuation, a CMRR figure of 70 dB is enough to reach the 100 dB target. Highly mismatched electrode impedances would result in rejection values below this limit, because of the potential divider effect. The TLC2202 input impedance is not listed on its datasheet, but because of the observed CMRR degradation it can be estimated to be dominated by an input capacitance of roughly 30 pF ", " Flexor mechanism coupled to the actuator. For the rotation mechanism, the actuator was coupled to the rotation axis of the stem holding the flexor mechanism actuator; the slide is preserved as a guide for curvilinear motion that will have the mechanism (Fig. 7). Fig. 7. Rotation mechanism coupled to the actuator. Additionally, supporting pieces were designed to allow mechanisms adjustment according to the morphology of the patient's hand, so that the characteristics of the patient\u2019s hand are non-limiting for use, Fig. 8. IV. SIMULATION OF THE EXOSKELETON MECHANISM SimMechanics MatLab software module to simulate the kinematic behavior of the mechanisms from control laws for the trajectory tracking designed, it was used. It provides a simulation environment for 3D mechanical multibody systems, from the models designed in SOLIDWORKS. Each of the models was translated by the program as a system block diagrams, which showed each component and junction that contains the mechanism involved in the movement. The flexor mechanism is shown in Fig", " Refrigerator temperatura \u2013 samples and reagents \u2013 pharmaceutical quality control lab. Optimal range between white lines. Fig. 5. Incubator \u2013 pharmaceutical quality control lab. Looking at the equipment power signals, the ratio of the total harmonic distortion is above the values approved by the standards and regulations. In the case of the Fig. 7, the first box shows the sinusoidal signal that the equipment (a hematological counter) receives, and the second box shows the analysis of such signal with a THD ratio of 52.96%. As evidenced in Fig. 8, this impacted the measurement of red cells and plates. Moreover, Fig. 9 shows a uniform sinusoidal signal, although it exceeds the limit prescribed by the regulation, being of 16.005%. 694 V. Gonzalez and S. Lorandi IFMBE Proceedings Vol. 49 IV. CONCLUSIONS The designed monitoring system verified that the centers under analysis do not comply with the regulatory standards applicable to the storage temperature for samples and reagents, water conductivity, and electrical noise. The impact to the laboratory test results is random, being impossible to confirm the universality of the error", " TAMARINDO (CHALEN) 29 20 TOTAL BY SEX 638 522 As seen from the Fig. 7 and consistent with community lifestyle main health problems correspond to digestive and respiratory illnesses and those are the main cause of consultation. Figure 7. Prevailing Disieses as reported thus far in the EFHR. IFMBE Proceedings Vol. 49 0.00% 5.00% 10.00% 15.00% 20.00% 25.00% 30.00% 35.00% 40.00% 45.00% 50.00% 0 - 1 2 - 9 10 - 19 20-64 Over 65 Underweight. Sev. Underweight. Mod Underweight. Acceptable. Normal Overweight Preobese Obese Type I Obese Type II Obese Type III Figure 8. Nutritional profile for population grouped bya age (% of population group). Other significant information is the nutritional profile as seen in Fig. 8. Nutritional profile was obtained by the use of body/mass index. The percentage of individuals at each age group was obtained by dividing the number in each classification by the total number in the particular age group. Notice the significant amount of underweight for the 0 \u2013 1 and the 2 \u2013 9 age groups. This is transcendental since it clearly shows chronic nutrition deficit translating into poor cognitive and physical development. The project provides a specifically applied object oriented Web software, pursuant to rules adopted by the MPH of Ecuador", " (6) shows areas defined ber where th ntire surface ttenuation measu to be consid radiographic adiographic para Va 80 500 150 300 f an apron in the a measureme as gonadal e collimated of the detecto rement technique ered for inspe technique f meters lue Units KV milise mAS Ma ir gonads region nt on a lead and shown beam reache r. ction of atten actors and d c. apron, radiati the ionizati s just to cov uis- ng on er Fig. 7 Image of the ionization chamber Bu sh Figure show cky as a ref ould complet s foreground erence for adj ely cover the remote ioniza usting the lig detector. tion chamber ht collimator with that IFMBE Proceedings Vol. 49 gi go im gr an Fig. 8 Measuremen ons shown in nads and the portant to pro IV A low perce ity (5%), ver d testing with Regions defined ts made on Figure (8) co region of t tect with apr . RESULTS OF ntage of ERP ified by the v fluoroscopic as the mediastinu aprons were rresponding t he mediastin ons. MEASUREMEN was detected isual aspect / image.. m and the gonad located in th o the zones o um that are TS with altered haptic inspe s e ref the most intection A and cont irreg W with as d feren 99.2 60% T twee twee tecti R/m V tion irreg and also or fl and store brea T 95% or fr bette that ing w It its l have instr atten Is with altho able prons with lo fractures sho act with the h ularities in ra ith the dosim out attenuatio irect radiatio t studied ran % attenuation attenuation, he 20 thyroid n 94% and 9 n 96", "45 mm Figure 7 Foot average measures for Mexican users 0 10 20 30 40 50 60 70 80 90 100 -100 0 100 200 300 400 500 600 700 800 Fuerzas en PF1 % Ciclo de la Marcha F R S e n N 0 10 20 30 40 50 60 70 80 90 100 -30 -25 -20 -15 -10 -5 0 5 10 15 Angulo de la articulacion tobillo-pie % Ciclo de la Marcha A n g u lo a rt ic u la c io n G ra d o s 0 0.5 1 1.5 2 2.5 3 -200 0 200 400 600 800 1000 1200 Momento en articulacion tobillo Derecho Tiempo M om en to e n ar tic ul ac i\u00f3 n (N -m m ) IFMBE Proceedings Vol. 49 Therefore, the Solidworks\u00ae software is used in order to stablish parameters for the model to make any necessary modification based upon stablished needs. Figure 8 Impulsion system constituted by three carbon fiber springs III. RESULTS It has been defined the ideal geometry for each of the prosthetic components especially for the impulsion system constituted by three carbon fiber springs as well as for a small system to replicate the movements of inversion and eversion located among the superior and the inferior plaques of carbon fiber. According to the above, the three arcs that form the arciform structures of a human foot such as height of the subastragalar joint are capable of supporting a 80 kg pacient" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002749_s11071-020-05932-9-Figure10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002749_s11071-020-05932-9-Figure10-1.png", "caption": "Fig. 10 Self-built 3D pendulum platform", "texts": [ " In summary, it can be concluded from the simulation results that the proposed robust adaptive finite-time attitude controller is capability of achieving the highperformance attitude tracking control of a 3D pendulum in finite time even in the presence of external disturbance. In this section, the hardware experiments are carried out to further illustrate the proposed robust adaptive finite-time attitude controller. A self-built 3D pendulum platform is utilized for the hardware experiments as shown in Fig. 10. The rotational motion of the 3D pendulum platform is controlled by three reaction wheels and three fans along the principal axes of inertia. The fans are used to provide the additional control torques when the speeds of the reaction wheels are saturated. Both the reaction wheels and the fans are actuated by the AC servo motors. Furthermore, the attitude and angular velocity of the 3D pendulum can be measured by the sensors and encoders installed in the servo motors. The servo motors are connected with the host computer, in which a motion control board is utilized to transform the sensor signals and generate the control commands" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002131_ecc.2016.7810331-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002131_ecc.2016.7810331-Figure2-1.png", "caption": "Fig. 2: Schematic of Quadrotor Manipulation System with relevant frames", "texts": [ " This paper is organized as follows: In section II, the considered robotic system is described, and kinematic and dynamic analysis are reviewed. The control problem to solve is formulated and the DOb and MPC approaches are described in section III. In section IV, simulation results using MATLAB/SIMULINK are presented. Finally, the main contributions are concluded in section V. 3D CAD model of the proposed system is shown in Fig. 1. The system consists mainly of two parts; the quadrotor and the manipulator. Fig. 2 presents a sketch of the proposed system with the relevant frames which indicates the unique topology that permits the end-effector to achieve arbitrary pose. The frames satisfy the Denavit-Hartenberg (DH) convention. The manipulator has two revolute joints. The axis of the first revolute joint (z0), that is fixed to the quadrotor, is parallel to the body x-axis of the quadrotor (see Fig. 2). The axis of the second joint (z1) is perpendicular to the axis of the first joint and will be parallel to the body y-axis of quadrotor at home (extended) configuration. Thus, the pitching and rolling rotation of the end-effector is now possible independently from the horizontal motion of the quadrotor. Hence, with this new system, the capability of manipulating objects with arbitrary location and orientation is achieved. By this non-redundant system, the end-effector can achieve 6- DOF motion with minimum number of actuators/links which is an important factor in flight", " The arm components are designed, selected, purchased and assembled such that the total weight of the arm is 200g, has maximum reach in the range between 22cm to 25cm, and can carry a payload of 200g. Three DC motors, (HS-422 (Max torque = 0.4N.m) for gripper, HS-5485HB (Max torque = 0.7N.m) for joint 1, and HS-422 (Max torque = 0.4N.m) for joint 2), are used. Each rotor j has angular velocity \u2126j and it produces thrust force Fj and drag moment Mj which are given by: Fj = Kfj\u2126 2 j (1) Mj = Kmj\u2126 2 j (2) where Kfj and Kmj are the thrust and drag coefficients. Let \u03a3b, Ob- xb yb zb, denotes the vehicle body-fixed reference frame with origin at the quadrotor center of mass, see Fig. 2. Its position with respect to the world-fixed inertial reference frame, \u03a3, O- x y z, is given by the (3x1) vector pb = [x, y, z]T , while its orientation is given by the rotation matrix Rb: Rb(\u03a6b) = C(\u03c8)C(\u03b8) S(\u03c8)S(\u03b8)C(\u03c8)\u2212 S(\u03c8)C(\u03c6) S(\u03c8)S(\u03c6) + C(\u03c8)S(\u03b8)C(\u03c6) S(\u03c8)C(\u03b8) C(\u03c8)C(\u03c6) + S(\u03c8)S(\u03b8)S(\u03c6) S(\u03c8)S(\u03b8)C(\u03c6)\u2212 C(\u03c8)S(\u03c6) \u2212S(\u03b8) C(\u03b8)S(\u03c6) C(\u03b8)C(\u03c6) , (3) where \u03a6b=[\u03c8, \u03b8, \u03c6]T is the triple ZY X yaw-pitch-roll angles. Note that C(.) and S(.) are short notations for cos(.) and sin(.). Let us consider the frame \u03a3e, O2- x2 y2 z2, attached to the end-effector of the manipulator, see Fig. 2. Thus, the position of \u03a3e with respect to \u03a3 is given by pe = pb +Rbp b eb, (4) where the vector pbeb describes the position of \u03a3e with respect to \u03a3b expressed in \u03a3b. The orientation of \u03a3e can be defined by the rotation matrix Re = RbR b e, (5) where Rbe describes the orientation of \u03a3e w.r.t \u03a3b. In [4], the equations of motion of the proposed robot have been derived in details. The dynamical model of the quadrotor-manipulator system can be written as follows: M(q)q\u0308 + C(q, q\u0307)q\u0307 +G(q) + dex = \u03c4 ; \u03c4 = Bu (6) where q = [x, y, z, \u03c8, \u03b8, \u03c6, \u03b81, \u03b82]T \u2208 R8 is the vector of the generalized coordinates, M \u2208 R8\u00d78 represents the symmetric and positive definite inertia matrix of the combined system, C \u2208 R8\u00d78 is the matrix of Coriolis and centrifugal terms, G \u2208 R8 is the vector of gravity terms, dex \u2208 R8 is the vector of the external disturbances, \u03c4 \u2208 R8 is the vector of the generalized input torques/forces, u = [F1, F2, F3, F4, \u03c4m1 , \u03c4m2 ]T is the vector of the actuator inputs, B = HN is the input matrix which is used to generate the body forces and moments from the actuator inputs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003652_j.addma.2021.102095-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003652_j.addma.2021.102095-Figure1-1.png", "caption": "Fig. 1. Flat specimen for mechanical investigations, a) build orientation and part location of additively manufactured ZYX specimens in the build space according to ASTM F2971, b) slicing form for different layer heights in a cross sectional view through the zx-plane. (For interpretation of the references to colour in this figure, the reader is referred to the web version of this article.)", "texts": [ " The delivery state of filament is on vacuum-packed spools with 2.85 \u00b1 0.10 mm in diameter. The fiber weight content is approx. 12.5 wt%, the fiber diameter is 7 \u00b5m and the fiber length distribution is 150\u2013400 \u00b5m. The extrusion-based additive manufacturing was done by a commercial system (Ultimaker 2 + Extended, Ultimaker) with the slicing software Simplify3D. Selected manufacturing parameters are given in Table 1. The part orientation in the built space showing the macroscopic waviness is illustrated in Fig. 1(a). In addition, Fig. 1(b) clarifies the differences in slicing due to the change in layer height. Thereby, the blue color P. Striemann et al. Additive Manufacturing 46 (2021) 102095 indicates the outer perimeter and the green color the inner 100% infill. The specimens were tested with both untreated as built and polished surface textures. For the polished surfaces, the process-related waviness is post-processed by stepwise polishing with different grain sizes (P 180 \u2013 P 320 \u2013 P 600 \u2013 P 1000). Finishing of the radius was carried out by using a grinding device with predetermined final contour" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003111_978-3-319-26500-1-Figure5.12-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003111_978-3-319-26500-1-Figure5.12-1.png", "caption": "Fig. 5.12 Impact of the focal length on the appearance of a model (piston rod) in the image. a f = 1 mm, distance = 75 mm b f = 4 mm, distance = 300 mm c f = 12.5 mm, distance = 900 mm d f = 32 mm, distance = 2400 mm", "texts": [ "8 Computation of the distance do of the object to the focal point f of the camera using the intercept theorem . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 80 Figure 5.9 Frustum of a normal map, zero-padded to match the volume sizes of model and frustum. . . . . . . . . . . . . . . . 82 Figure 5.10 Voxelized model, zero-padded to match the volume sizes of model and frustum . . . . . . . . . . . . . . . . . . . . . . . . 82 Figure 5.11 Correlation function of the two volumes of Figs. 5.9 and 5.10 . . . . . . . . . . . . . . . . . . . . . . . . . . . . 83 xii List of Figures Figure 5.12 Impact of the focal length on the appearance of a model (piston rod) in the image . . . . . . . . . . . . . . . . . 85 Figure 5.13 Analysis of the accuracy of the orientation estimation technique dependent on the focal length of the camera . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 87 Figure 5.14 Analysis of the accuracy of the orientation estimation technique dependent on the noise level present in the normal map . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ", "5 Note that the 1/3 inch sensor of the camera has a crop factor of 8, which means that the 35 mm equivalent focal lengths are between a fish-eye (8 mm) and a super telephoto lens (512 mm). At the short end the models are significantly distorted but the perspective scaling is large. At the long end the viewing rays are nearly parallel, which results in nearly zero perspective scaling and distortion. All scenes were rendered such that the objects approximately filled 10 % of the frame. As the focal length obviously changes the field of view, the distance between the camera and the object varied from 75 at 1 mm focal length to 4800 at 64 mm focal length (see Fig. 5.12). The orientation estimation is the first of the two steps performed for pose estimation. Based on the result, the translation is estimated. Therefore, it is essential that the orientation estimation is robust and accurate. To analyze the accuracy, different 5The simulated focal lengths were 1, 2, 4, 8, 12.5, 16, 32 and 64 mm. 86 5 Normal Map Based Pose Estimation models and different parameters were analyzed for their impact on its results. The most important aspect is the focal length of the camera" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001829_icra.2016.7487631-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001829_icra.2016.7487631-Figure3-1.png", "caption": "Fig. 3: This figure illustrates the relation between the youbot base and the sensor reference frame. The relative pose of the sensor, with respect to the robot frame, is ks. While the mobile base moves along the trajectory oi, the sensor will move along the path zi in its own reference frame. As expressed by the kinematics of the robot, kbb is given by the average between the two baseline, i.e kbb = (kb1+kb2)/2", "texts": [ " ui = ( tli t r i ) consists of the tick increments reported by the left and the right wheels, denoted respectively as tli and tri . The direct kinematics f(k,ui) can be approximated as: oi = f(ko,ui) = (krtri+kltli) 2 0 krtri\u2212kltli 2kb (14) 2) Omnidirectional Robot: The odometry parameters are ko = ( kfl kfr kbl kbr kbb )T , (15) where the first four parameters are the coefficients for respectively front left, front right, back left and back right wheels. kbb is a coefficient defined as the average between baseline of the wheels and the distance between the front and rear axes, as shown in Figure 3. The encoder ticks are stored in a 4 dimensional vector uT i = (tfli t fr i t bl i tbr i )T as oi = \u22121/4 1/4 \u22121/4 1/4 1/4 1/4 \u22121/4 \u22121/4 1/4kbb 1/4kbb 1/4kbb 1/4kbb kfltfli kfrtfri kbltbl i kbrtbr i (16) IV. ISSUES OF CALIBRATION USING LEAST SQUARES Most of the calibration procedures in the literature rely on least squares minimization. The general procedure consists in moving the robot along a predefined trajectory, while recording its encoder ticks. During the acquisition, a ground truth of the position of the mobile base or its sensors is obtained through some external observer or by algorithms that process only the exteroceptive sensor input" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002485_13645706.2020.1755313-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002485_13645706.2020.1755313-Figure2-1.png", "caption": "Figure 2. Schema of the tip of the forceps fitted with a microelectromechanical system (MEMS) triaxial force sensor. The sensor was 14mm in length, 7mm in width and 2mm in height. This device could measure the grip force in the X, Y and Z directions.", "texts": [ " The system was composed of grasping forceps (Karl Storz, Tuttlingen, Germany, KK33322CC), a microelectromechanical system (MEMS) triaxial force sensor attached to the tip of the forceps, a bridge and amplifier circuit board attached to the handle of the forceps, a microcomputer unit (MCU) board capable of analog to digital conversion (ADC) and serial communication, and a laptop PC connected to the MCU board by a USB cable (Figure 1). To mount the MEMS triaxial force sensor [2] at the tip of the forceps, we reduced the sensor size to 14mm in length, 7mm in width, and 2mm in height (Figure 2). The force-sensitive area was approximately 24mm2 and a thermistor was included for temperature compensation. The magnification of the amplifier circuit driven at 3.3V was 256 times for vertical force and 192 times for shear force. The ADC device had a 10-bit resolution, and the triaxial force and temperature data could be transferred with serial communication at a rate of up to 250Hz. The measurement range, force resolution and sensitivity of the fabricated MEMS sensor were 5N, 17mN and 0.45V/N for vertical force and \u00b11" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000672_robio49542.2019.8961446-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000672_robio49542.2019.8961446-Figure2-1.png", "caption": "Fig. 2. Robot performs constrained task on curved surface", "texts": [ " min( ) a x x min( ) b x x Normalized each vector: / 2i ik a b N Cosine similarity: ( ) ( ) cos( ) k k k k a b a b 7. 2 min(cos( ) 1) fProduce new by PSO 8. end 9. DMP_ reproducing (): [ , , ]dmp dmp dmpx x x Considering an unknown surface, we use force control to make the end effector of the arm keep contact with the surface to ensure that the generalized trajectory lies on the surface. We directly add the acceleration term to initial DMP and consider six-dimension force control. Fig. 2 shows robot performing a constrained task on the curved surface. In order to keep contact with the surface, the robot must maintain a certain contact force on the surface. In order to keep alignment with the normal vector of the surface, the robot must adjust the orientation of the end effector to be perpendicular to the surface. It means the contact force must be controlled to a given force. The force error is defined as [14]: T T = 0 0 0 0 0T T zd T T T T T T x y z x y z F F F F M M M F (10) where T zdF is the desired contact force in the Z direction represented in coordinate system {T} , T xF , T yF , T zF , T xM , T xM , and T xM are actual force and torque measured by six-axis force sensor" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001820_0954408916652648-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001820_0954408916652648-Figure7-1.png", "caption": "Figure 7. Finite element model of the spindle\u2013toolholder", "texts": [ " Based on the slope and intercept of the at UNIV OF VIRGINIA on June 7, 2016pie.sagepub.comDownloaded from fitting function, the fractal dimension and fractal roughness parameter can be obtained as Ds \u00bc 1:36 and G \u00bc 6:238 10 12. The three-dimensional fractal dimension D can be written as D \u00bc Ds \u00fe 1 \u00bc 2:36. The taper surface of spindle\u2013toolholder joint is assumed flat in macro-scale and the contact pressure of distribution can be obtained from the static analysis of spindle\u2013toolholder assembly in the ANSYS software, as shown in Figure 7. The node number of taper surface is nT \u00bc 120. The CONTAC174 and TAGET170 units are used to construct the contact pairs of joint surface, in which the taper surface of toolholder is set as the contact surface and that of spindle is set as target surface. The draw-bar force is set as 8 kN. The contact pressure of each node in the taper surface can be extracted from the general Postproc section of ANSYS software. The contact pressure distribution of the taper surface along the axis of toolholder is shown in Figure 8, in which the origin of coordinate represents the point with the largest dimension of taper surface" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003279_j.matpr.2020.12.094-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003279_j.matpr.2020.12.094-Figure2-1.png", "caption": "Fig. 2. Physical 3D dummy model created in ANSYS R18.1 workbench.", "texts": [ " Restricted segment examination (FEA) is a motorized strategy for anticipating how a system responds honestly to vibration, powers, temperature, fluid stream, and other physical repercussions. Restricted part examination displays whether a system will damage, decimate, or toil the way it was described. It is named as examination, however used to predict what will happen when the system is used. A 3D dummy model having a dimension of 40 mm 40 mm 2 mm rectangular work-piece was generated to scrutinize the laser beam process during simulation as shown in Fig. 2. The hole geometrical feature was created on this rectangular work-piece followed by adequate mesh refinement to achieve proper results during ES-LBTM approach. The properties of DSS super alloy and the conditions for machining environment which were chosen during this research is mentioned in Table 1 and 2 respectively. The optimized values those are obtained from PSO single objective optimization (from section 2.1) were adopted as a boundary condition for further analysis. As from the preliminary experimentation it has been observed that the laser power was varied; which influences greatly out from the other parameters so the laser power was chosen as variable in ANSYS simulation, and boundary conditions during the analysis were considered as follows" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003662_s40684-021-00357-0-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003662_s40684-021-00357-0-Figure3-1.png", "caption": "Fig. 3 Mesh model of the L-PBF printer, and refined meshes adopted around the workbench and along the laser scanning path", "texts": [ " The first part of the right side of this equation represents drag force, which is a function of the relative velocity, the second part is gravity force, and the third part is additional forces, including Saffman lift and others (user defined). (5)q = 2AP R2 e \u22122r2 R2 (6) du p i dt = FD ( ui \u2212 u p i ) + i ( p \u2212 ) p + Fi p Considering the complex geometry of the L-PBF printer, the computational domain is divided into many sub-domains and the mesh types are set reasonably to improve the quality of the mesh. This mainly increases the mesh density around the laser scanning path, inlets, the outlet, and pipeline walls. As shown in Fig.\u00a03, the diameter of the laser spot is 200\u00a0\u03bcm. To accurately simulate the process of laser scanning, a crossscale mesh from the micrometer-to-millimeter scale is generated in the middle of the workbench. The gas flow near the pipeline wall is more complex. Thus, the mesh density is increased at the pipeline boundary. A boundary layer with five layers and a 1.2-fold growth rate is used to increase the mesh density and thus ensure accurate calculations. To verify mesh independence, 18 points (A1\u2013C3\u2032) evenly distributed on the two planes above the workbench are selected, as shown in Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000019_iros.2018.8593927-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000019_iros.2018.8593927-Figure1-1.png", "caption": "Fig. 1: Dual-arm robot with closed kinematic chain.", "texts": [ " The environment can be represented by a threedimensional voxel grid and its Euclidean distance transform (EDT). Thus, the distance computation between a sphere and obstacles is reduced to a look-up operation in the distance map. Similar to CHOMP, this second approach is used in this paper, because the complexity depends solely on the number of spheres and the resolution of the discretized workspace, and not on the number of obstacles. See [20] for details. The equality constraint (1g) enforces the closed kinematic chain, that is illustrated in Figure 1 for a dual-arm robot. It can be defined in terms of the transformation matrices of the left arm T L(q) and of the right arm TR(q) as T L(q) = T rel TR(q) (5) whereby the relative transformation T rel describes the desired difference in position and orientation between the two end-effectors. In order to reduce the number of constraints, translation and rotation are constrained separately. For simplicity, the position vectors are notated as tL(q) = Position{T L(q)} (6) tR(q) = Position{T relTR(q)} (7) and the orientations as four-dimensional quaternions QL(q) = Quaternion{T L(q)} (8) QR(q) = Quaternion{T relTR(q)} " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001047_icems.2014.7013792-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001047_icems.2014.7013792-Figure4-1.png", "caption": "Fig. 4 Coordinate transformation of flux linkage for three-phase SRM", "texts": [ " For the purpose, when the flux is located in the zone k (k=0, \u2026, 5), Vk+1 and Vk+2 are selected to increase the torque amplitude since they result in acceleration of the flux in the direction of the rotor speed; Vk-1 and Vk-2 are chosen to decrease the torque amplitude; Vk+1 and Vk-1 are chosen to increase the flux amplitude; Vk+2 and Vk-2 are chosen to decrease the flux amplitude. As stated above, a switch table is summarized in Table 1, suitable for executing space vector control of three-phase switched reluctance motor after given the analysis of flux linkage and torque regulation rules. In order to express the stator flux linkage vector, it is necessary to transform the three independent flux vectors into an orthogonal - coordinate as is shown in Fig. 3. From Fig.4, the following formulas can be easily obtained by coordinate transformation, \u2212= \u2212\u2212= 60sin60sin 60cos60cos CB CBA \u03c8\u03c8\u03c8 \u03c8\u03c8\u03c8\u03c8 \u03b2 \u03b1 (2) The amplitude and vector angle of stator flux linkage are defined as below: = += \u03b1 \u03b2 \u03b2\u03b1 \u03c8 \u03c8 \u03b8 \u03c8\u03c8\u03c8 arctan 22 S (3) According to the theory of VSVC of switched reluctance motor, we have to know the space position of the compound flux linkage before we judge which vector should be selected. The angle formed between the compound flux linkage vector and axis can be calculated using the formula (3) and then the interval for the compound flux linkage in the stator coordinate system is also established" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003349_s00170-021-06813-0-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003349_s00170-021-06813-0-Figure15-1.png", "caption": "Fig. 15 Exploded CAD model of cold plate including BN tubes, gauge 20 copper plate (0.8-mm thick), encapsulated by conventional manufacturing of three-piece SLS aluminum alloy block (final envelope size is 100 mm \u00d7 185 mm \u00d7 18 mm)", "texts": [], "surrounding_texts": [ "Active cooling prototype is represented by following equations that govern incompressible fluid flow of LM and heat transfer: The continuity equation (conservation of mass) \u2207:V \u00bc 0 \u00f04\u00de The momentum equation (Navier\u2013Stokes equations) \u03c1 \u2202V \u2202t \u00fe V :\u2207V \u00bc \u2212\u2207p\u00fe \u2207:\u03c4\u2212\u03c1g \u00f05\u00de Conservation of energy (the temperature distribution): \u03c1Cp \u2202T \u2202t \u00fe V :\u2207T \u00bc \u2207: k\u2207T\u00f0 \u00de \u00fe Q \u00f06\u00de where V is velocity of LM, \u03c1 is density of LM, T is temperature, p is pressure, t is cycle time, \u03c4 is deviatoric stress, g is gravity vector, \u03b2 is volumetric expansion coefficient, Cp specific heat at constant pressure, k thermal conductivity, Q is volumetric heat source. Ansys Fluent is used for computational fluid dynamics (CFD) simulation to provide approximate solutions to nonlinear partial differential governing equations. The equations are therefore converted into a set of algebraic equations solved through numerical techniques. Preprocessing efforts include adaptive tetrahedron mesh (Fig. 10), assigning material properties, and full contact is assumed at interface of plated tube and SLS aluminum alloy metallic block (100\u00d7185\u00d718mm3). Due to the limitation of the CFD software available by today\u2019s simulation methods, it only allows \u201cfull contact\u201d at interfaces, and consideration of thermal contact resistance is not feasible. Specific boundary condition available in Ansys (fan boundary condition) is used to split the tube as shown in post-processing (Fig. 11) to apply the characteristics curve of the diaphragm pump presenting flow rate vs pressure for greater fidelity of LM flow model. In solving the energy equation resulting temperature contours, the conjugate heat transfer is used to describe heat transfer that involves variations of temperature within solids and liquid metal due to thermal interaction between the solids and LM as seen in Fig. 12. In this simulation, surface heat is generated from mounted electronics (72 W), and heat is taken away by the AM cold plate and free convection. Heat flux post-processing report has shown that heat transferred by free convection is less than 10% and most of heat dissipated to the cold boundary condition via conduction. CFD simulations performed for in situ AM cold plates including plated ceramic tubes replicating test runs and the pre- and post-processing corresponding to in situ AM cold plate of experiment run #1 (BN-PC-IAM) are shown in Figs. 11 and 12. To check the convergence of simulation, residual curves of continuity, velocity, energy, and user-defined scalar (uds-0) are plotted. As seen in Fig. 13 after 950 iterations, the numerical model was fully converged concluding validation of postprocessing results and appropriate mesh grid size without any need for further refinement. The CFD simulation revealed key characteristics and behaviors regarding the design of the experiment including diameter of tube, thickness of block, and selection of appropriate pumping for required flow rate. The outcome of iterative simulations resulted in optimized prototypes to cool and meet the temperature limit of mounted electronics. However, verification and validation of models are required before committing to performance of developed cooling prototypes. In the following section, actual thermal test is performed for fabricated prototypes. Post-processing of simulations and experimental results are compared, and CFD simulation is with an error margin below 6% compared to test results. The comparison is demonstrated at \u201cThermal test process and experimental results\u201d section (Table 2). Fig. 16 Fabricated cold plates by in situ AM and conventional techniques (similar geometry and material for pair comparison). a In situ SLS cold plate including copper-plated BN ceramic tubes creating superior bonding by fusion of copper with aluminum alloy AM powder at thermal interfaces. b In situ SLS cold plate including silver-plated BN ceramic tubes and fusion of silver with aluminum alloy AM powder has shown highest efficiency. c Conventionally assembled cold plate including three-piece SLS aluminum alloy block, gauge 20 copper off-the-shelf plate. d Conventionally assembled cold plate including three-piece SLS aluminum alloy block, gauge 20 silver off-the-shelf plate. Three-piece assembly fastened with #4-40 SS screws torqued to 4.7 lb-in. surface flatness .003 \u201c(0.08 mm) and surface roughness 64 \u03bc-in. (0.002 mm) at top surface (interface to heat source) and bottom surface (interface to Peltier cooler) for all pieces" ] }, { "image_filename": "designv11_22_0002985_s12206-020-1117-0-Figure16-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002985_s12206-020-1117-0-Figure16-1.png", "caption": "Fig. 16. Contact stress of internal gear pair with two contact points (unit: MPa).", "texts": [ " '' '' '' z x y z t t t t t t t t t t t n n n n l t t \u03b4 \u03b4 \u03b4 \u03b4 \u03d5 \u03d5 \u03b1 \u03b1 \u03d5 \u23a7 \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23aa \u23a8 \u23aa \u23aa \u23aa \u0394\u23aa \u23aa \u0394 + \u0394 + \u0394 \u23aa \u23aa + \u23aa \u2202\u03a9 \u2202\u03a9 \u2202\u03a9\u23aa\u0394 = \u00d7 \u22c5 =\u23aa \u2202 \u2202 \u2202\u23a9 \uff08 \uff09 \uff08 \uff09 \uff08 \uff09 \uff08 \uff09 (31) Based on solid modeling methods in Sec. 3, the internal gear models with two contact points can be established as shown in Fig. 14. For its stress analysis, the contact unit size, material properties, boundary conditions and load application are the same with gears with single contact point. The numbers of units and nodes are 378644 and 1608008, respectively. Finite element model of internal gears with two contact points is shown in Fig. 15. The analysis results in Fig. 16 show that the maximum contact stress of internal gears with two contact points is 616.88 MPa. The maximum von Mises stress and shear stress of the pinion with two contact points in Fig. 17 are 417.21 MPa and 133.82 MPa, respectively. The maximum von Mises stress and shear stress of the internal gear with two contact points in Fig. 18 are 347.72 MPa and 165.97 MPa, respectively. Obviously, two peak stress regions can be found on tooth profiles, which are also corresponding to two contact points. The stress distribution area is relatively concentrated and it has the trends of expanding towards the tooth root direction" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003518_s11071-021-06407-1-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003518_s11071-021-06407-1-Figure3-1.png", "caption": "Fig. 3 Earth and vessel body fixed coordinates", "texts": [ " In Theorem 1 of this paper, the new nonlinear membership function v\u03b2i\u03b2ee(\u03b4 \u2032 e\u03b2k) is: v\u03b2i\u03b2ee ( \u03b4\u2032 e\u03b2k ) = ( \u03b4\u2032 e\u03b2k \u2212 \u03b4\u2032 \u03b2i\u03b2e=2k ) / ( \u03b4\u2032 \u03b2i\u03b2e=1k \u2212 \u03b4\u2032 \u03b2i\u03b2e=2k ) = ( \u03b4 \u03b2 ek \u2212 \u03b4 \u03b2 i\u03b2e=2k ) / ( \u03b4 \u03b2 i\u03b2e=1k \u2212 \u03b4 \u03b2 i\u03b2e=2k ) = v\u03b2i\u03b2ee ( \u03b4 \u03b2 ek ) (45) Thus v\u03b2i\u03b2ee(\u03b4 \u03b2 ek) is in nonlinear form for \u03b4 \u03b2 ek when \u03b2 > 1. In other words, the nonlinearities in stability conditions are extracted by v\u03b2i\u03b2ee(\u03b4 \u03b2 ek), and the resulting stability conditions as inTheorem1are transformed into linear form. A vessel dynamic positioning control system in [39] is used to verify the proposed method. In Fig. 3, a vessel body fixed coordinate xoy and an inertial coordinate NOE are established. The vessel position and heading in NOE coordinate is described as \u03b7 = [n e \u03c8]T \u2208 R3, the speed of xoy relative to NOE is defined as \u03bd = [u v r ]T \u2208 R3. And the fuzzy systemmatrices are given below: Ai = [ 03\u00d73 Ri 03\u00d73 \u2212G\u22121D ] , R1 = \u23a1 \u23a3 \u03b11 \u03b12 0 \u2212\u03b12 \u03b11 0 0 0 1 \u23a4 \u23a6 , R2 = \u23a1 \u23a3 1 0 0 0 1 0 0 0 1 \u23a4 \u23a6 , R3 = \u23a1 \u23a3 \u03b11 \u2212\u03b12 0 \u03b12 \u03b11 0 0 0 1 \u23a4 \u23a6 B1 = B2 = B3 = E1 = E2 = E3 = [ 03\u00d73 G\u22121 ] where the vessel heading x3 can be seen as the premise variable of the fuzzy system, and \u03b11 = cos(x3), \u03b12 = sin(x3), x3 = 1 3\u03c0 " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002520_j.matpr.2020.05.438-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002520_j.matpr.2020.05.438-Figure3-1.png", "caption": "Fig. 3. Cross axis wind turbine.", "texts": [ ", tip speed ratio (k), coefficient of power (Cp), and coefficient of torque (Ct) for varied wind conditions (Va) ranging from 4 m/s to 10 m/s. Table 2 Specifications of cross axis wind turbine [27]. Aerofoil profile NACA 0018 Diameter of rotor, D 35 cm Horizontal blade Pitch angle, bh 0 degree Chord length, C 3.40 cm Length, L 15 cm Number of blades 6 Vertical blade Chord length, c 5 cm Pitch angle, bv 0 degree Height, h 30 cm Number of blades 3 Cross axis wind turbine (CAWT) is a transformed version of VAWT having the characteristics of the typical VAWT type as well as HAWT type. As can be referred in Fig. 3, CAWT consist of 3 vertical blades as well as 6 untwisted horizontal blades. The vertical blades are interconnected to the horizontal ones through connectors. The horizontal blades play a role as a connecting strut linking the CAWT arrangement to the hub. The lower horizontal blade arrangement and the upper horizontal blade arrangement are positioned at an angle of 60 degree offset with respect to each other (refer Fig. 3). The horizontal blade profile is made using NACA0018. Similarly, the vertical blade profile is made using NACA0018. The dimensional details of both the horizontal and vertical blades as well as the CAWT are listed in Table 2. Fig. 4 illustrates the CAD (computer aided design) models of different CAWT parts. All the CAD models are developed using SOLIDWORKS and it is converted Table 1 Properties of ABS [10,26] Properties Values Tensile modulus 2.21 Gpa Flexural strength 75.84 Mpa Tensile strength 44", " Based on the earlier prepared code, the printer head lays the materials layer by layer. Thus, all the parts of CAWT are produced based on additive manufacturing concept. The parts produced by FDM method requires finishing touch. The printing surface is smoothened and irregularities if any can be removed by chemical treatment with acetone vapour, file finishing or using sand paper. The specification of additive manufacturing machine based on FDM technique available is listed in Table 3. With reference to Fig. 3, all the CAWT parts produced by additive manufacturing concept are assembled together. Fig. 5 shows the assembled CAWT and its dimensional detail [27] is listed in Table 2. Fig. 6 illustrates the experimental test setup erected at HITS, Padur, Tamilnadu to analyze the performance of CAWT. The test rig consists of an open jet wind tunnel and the rotational speed of the axial fan of the open jet wind tunnel can be varied from minimum to its maximum rated speeds. The variations in its rotational speed are done by using the ABBTM make variable frequency drive" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002384_s12206-020-0325-y-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002384_s12206-020-0325-y-Figure7-1.png", "caption": "Fig. 7. Design variables for FE shape optimization, viewed from an axial direction (unit: mm).", "texts": [ " 5, and their initial and final optimum values are summarized in Table 3. The initial shape design parameters, wup, wdown, r, d, l, and h in Fig. 5 are an outer width, an inner width, a fillet radius, a radial height to the protrusion valley, and the valley gap, respectively. The move limits of design variables, x1 to x8, are depicted in Figs. 6 and 7. The variable x1 is a moving limit of the protrusion top; x6 is a moving limit of a valley bottom; x7 and x8 are a horizontal move limit of the protrusion top as in Fig. 6. In Fig. 7 that shows one of the Table 1. Mechanical properties. same two inclined side surfaces, the variables x2 and x3 are move limits of the top edge and x4 and x5 are move limits of the bottom edge. The shape optimization of the composite protruded interface structure proceeds iteratively in order to refine the initial design (Fig. 5) through the shape variables (i.e., coordinates of FE nodes), based on the computed elemental stresses satisfying Eqs. (1)-(3) and by changing elemental shapes in the range of Eqs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000279_j.nahs.2019.06.003-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000279_j.nahs.2019.06.003-Figure2-1.png", "caption": "Fig. 2. Clocks \u03b8 \u03b4RDT (left) and \u03b8 \u03b4MDT (right) corresponding to a parameter \u03b4 with jumps at t\u03b41 , t \u03b4 2 , . . . as well as range and minimum dwell-time, respectively.", "texts": [ " (d) In parallel to [6,7,9], Theorem 9 can be extended to also cover uncertain parameters with range or minimum dwell-time based on replacing \u03b8 with clocks \u03b8 \u03b4RDT and \u03b8 \u03b4MDT, respectively, that depend on the location of the discontinuities of the present uncertainty \u03b4 but not on its values. In the case of range dwell-time, i.e., parameters \u03b4 with discontinuities at times t\u03b41, t \u03b4 2, . . . with t\u03b4k+1 \u2212 t\u03b4k \u2208 [Tmin, T ] for all k \u2208 N and some Tmin \u2208 (0, T ), such a clock \u03b8 \u03b4RDT satisfies \u03b8 \u03b4RDT(t) = t \u2212 t\u03b4k on [t\u03b4k , t \u03b4 k+1), k \u2208 N and is illustrated on the left hand side of Fig. 2. Robust stability against all uncertainties \u03b4 : [0,\u221e) \u2192 [a, b] with range dwell-time is then assured by feasibility of the LMIs (14) with the jump condition (14d) replaced by( 0 0 0 AJ )T X\u0302(0) ( 0 0 0 AJ ) \u2212 X\u0302(\u03c4 ) \u227a 0 for all \u03c4 \u2208 [Tmin, T ]. In the case of minimum dwell-time, i.e., parameters \u03b4 with discontinuities at times t\u03b41, t \u03b4 2, . . . with t\u03b4k+1\u2212t\u03b4k \u2208 [T ,\u221e) for all k \u2208 N, a corresponding clock \u03b8 \u03b4MDT satisfies \u03b8 \u03b4 MDT(t) = t\u2212t\u03b4k on [t\u03b4k , t \u03b4 k+T ) as well as \u03b8 \u03b4MDT(t) = T on [t\u03b4k+T , t\u03b4k+1) for k \u2208 N and is depicted on the right hand side of Fig. 2. Robust stability against all uncertainties \u03b4 : [0,\u221e) \u2192 [a, b] with minimum dwell-time is guaranteed by feasibility of the LMIs (14) and, in addition, L (( 0 X(T ) X(T ) 0 ) , Pp(M(T )), ( \u03c8(T )G(T ) \u03c8(T ) ) ss ) \u227a 0 (16a) as well as L (( 0 R(T ) R(T ) 0 ) ,M(T ) + M(T )T , \u03c8(T )ss ) \u227b 0. (16b) One can show that feasibility of (14) and (16) is equivalent to feasibility of (14) and X\u0307(T ) = 0 as well as R\u0307(T ) = 0; the latter is more suitable for controller synthesis than (16) as we will see in Section 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003114_tmag.2015.2438872-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003114_tmag.2015.2438872-Figure3-1.png", "caption": "Fig. 3. Experimental set for the magnetic material properties.", "texts": [ " The mechanical air gap between the rotor and the stator is 1 mm. The dc supply voltage is 310 Vdc, and the maximum current is 10 Apeak. The torques in the washing and dehydrating modes are 37 and 2 Nm, respectively, and the speeds are 46 and 1200 r/min. The specifications are shown in Table I. To measure the B\u2013H curve and the iron loss under the stress and strain conditions that occur due to the slinky-lamination process, a ring core from a stator yoke of the proposed motor was prepared by a wire cut. The test conditions of the core are shown in Fig. 3 and Table II. The test was carried out over several levels of excitation field and at several frequencies. Insulation tape was wound around the ring core for insulation purposes. For both the primary coil N1 and the secondary coil N2, 200 and 40 turns were used to guarantee the exciting of a high field (over 5 kA/m). The B\u2013H curve and the iron loss test were conducted as per IEC 60404-6. In this test, the H field is measured at the primary coil, and the B field is measured at the secondary coil" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003187_s1061934815070199-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003187_s1061934815070199-Figure1-1.png", "caption": "Fig. 1. Schematic representation of the coating of the ccc TMAPL column.", "texts": [ " To obtain relatively steady and reverse EOF values, the inner wall of the capillary column can be modified with the functionalized materials such as polyethyle neimine [20] and quaternary ammonium nanopartic ulates [21, 22], which form the new open tubular cap illary electrochromatogarphy mode for the determina tion of anions such as nitrites and nitrates. In our previous work [23], we prepared a new trim ethylamine aminated polychloromethyl styrene nano latex and TMAPL coated capillary column (Fig. 1), which has been successfully applied for trace bromate determination by FASS\u23afOT CEC method. Built on this success, we have in the present work crafted two major aims as follows: (i) to validate the practicability of the ccc TMAPL column; and (ii)to develop a sen sitive FASS\u23afOT CEC method for the determination of nitrites and nitrates in plasma and urine for healthy persons and RA patients. This method possesses prac tical applications in a wide array of biological, clinical and pathophysiology analyses", " Working standard solutions were pre pared by diluting the stock standard solutions with ultra pure water. Each determination was performed in triplicate. All experiments were performed on an Agilent CE system with a diode array detector (Agilent Technolo gies, Waldbronn, Germany). The apparent pH was measured using a pH meter (Shanghai Weiye Factory, Shanghai, China). The dimension of the bare fused silica capillaries (Yongnian Optical Fiber Corpora tion, Hebei Province, China) was of inner diameter 50 \u00b5m and length of 50 cm (41.5 cm to the detection window). The ccc TMAPL (Fig. 1) is of 50 \u00b5m i.d. and 50 cm length (41.5 cm to the detection window) coated with flushing TMAPL nano latex through the bare fused silica capillary. The preparation procedures of ccc TMAPL column were same as mentioned in our previous work [23] by introducing TMAPL nano particles into a treated bare fused silica capillary for three times, the nanoparticles were coated onto the inner wall surface of the capillary by dipole dipole force, hydrogen bonding and electrostatic interac tions. The ccc TMAPL columns were characterized by infrared spectrum and EOF values", " After thawing at room temperature, a 300 \u00b5L urine sample was deproteinized using the same procedures as the plasma sample preparation. The 50 \u00b5L of supernatant was diluted 50 times with ultra pure water (50 \u00b5L of super natant with 2.45 mL water) to prepare the final urine samples for FASS\u23afOT CEC injection. This study was approved by the Ethics Committee of Zhengzhou University, and the informed consent was obtained from all subjects (adults of 20\u201340 years old). RESULTS AND DISCUSSION OT CEC separation condition for nitrite and nitrate in urine and plasma. As indicated in Fig. 1, the inner wall surface of the ccc TMAPL column is covered with positive charges, and its EOF will reverse. On one hand, it avoids the introduction of EOF reverse addi tive (such as cetyltrimethyl ammonium chloride) in BGE, which is sometimes undesirable because of its influence on the UV, conductivity, and mass spectro metric detection of inorganic and/or organic anions. On the other hand, the TMAPL nano latex coating as chromatographic stationary phase acts on the ion exchange interaction with nitrites and nitrates, which leads to the baseline separation (Rs > 1" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000096_042019-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000096_042019-Figure2-1.png", "caption": "Figure 2. Surface profiler pads of theTB: a - parallel channel intercusion bevel; b - helical surface; c - Raleystep", "texts": [ " Thrust bearings (TB) are used for sensing the axial load acting on the rotor, and transmit it to the stator, as well as to lock the rotor axially relative to the housing. When operating in the thrust bearing between the rotating disc 1 (Fig. 1) and pads 2 are formed thin lubricating layer 3, and in the angular direction are separated by aintercushion channels 4. These channels provide a supply of liquid lubricant in the layer. When rotating thrust disk in the channel 4 on the disc surface is formed a thin boundary layer 5 with a notional boundary 6 cooperating with lubricating layers adjacent cushions and fresh fluid channel 4. In Fig. 2 shows variations of the basic profiles of the carrier surface flat-wedge pads used in turbochargers [1]. The conducted experimental studies of the thrust bearing shown with fixed pads shown [2] that the surging of the centrifugal compressor (CC) [3] of the lubricating layer pressure waveform is a substantially non-stationary and occurs with changes in time before and during compressor pat (Fig. 3). The sensors D1 and D2 located at the average radius of the opposing pads on the working side of the bearing", " The dependence of the viscosity and density of the lubricant temperature: 2exp ( ) ( ) ,0 0 1 0t t t t (35) exp ( ) ,0 0t toil (36) where 000 ,, t temperature, viscosity, oil density on giving a intercushion channel; t- oil temperature; 1, - temperature coefficient of viscosity, depending on the grade of oil and the temperature range under consideration; oil - temperature coefficient of volumetric expansion of the oil. 7. The clearance in the thrust bearing with fixed pads. 7.1. The surface with parallel bevel intercushion channel (Fig. 2a) sin (1 ) cos2 . . 2 2 ( , , ) ( , , ) . 0 r d h h \u0443 ro d disp tilt \u043a R\u043a Hd T r \u0443 T r \u0443 dyp p p p p pp (37) 7.2. Helical bevel wedge surface (Fig. 2b) 1 cos2 . . 2 2 ( , , ) ( , , ) . 0 d h h \u0443 ro d disp tilt \u043a R\u043a Hd T r \u0443 T r \u0443 dyp p p p p pp (38) 7.3. Raley stepped surface (Fig. 2c) 20 . . 2 0 cos 2 ( , , ) ( , , ) . d d d disp step \u043a H p p p p p p p h h \u0443 r R T r \u0443 T r \u0443 dy (39) where 2h \u043e the thickness of the lubricant layer at the position of the disc at the origin; . .\u0443d disp thrust disk location coordinate is given its own equation or determined by solving the equations of the dynamics of the rotor (1); ( )1 2h htilt - the depth of the bevel at the front edge of the pad; MMBVPA IOP Conf. Series: Journal of Physics: Conf. Series 1158 (2019) 042019 IOP Publishing doi:10" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003486_j.jmps.2021.104411-Figure21-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003486_j.jmps.2021.104411-Figure21-1.png", "caption": "Fig. 21. Stress-driven test of one bi-stable joint of Prototype A (\ud835\udc5b = 2, \ud835\udc410 = 4.7 N and \ud835\udc3e = 1.80N/mm): (a) layout, with indication of the lever arm for the hanging weight with respect to the pitch point; (b) measured \ud835\udc40 vs. \ud835\udee5?\u0304? graph in loading\u2013unloading paths, compared with the theoretical predictions.", "texts": [ " In any case, non-locality does influence the energetic barrier that separates one inhomogeneous configuration from any other energetically equivalent to it because, in order to pass from one configuration to the other, the cable has to be further strained. The detailed dynamic characterization of the sequence of snaps will be the subject of a successive work. \ud835\udc41 4 a 4 t .2. Prototype A - Soft device Prototype A is now tested in a soft device. Only the results under stress-driven testing on one single joint and the shorter prototype re recorded, since the longer chain does not add any conceptual difference. .2.1. One single joint The two coupled segments, of length 33.8mm, lay in the vertical plane, as indicated in Fig. 21(a). The initial prestress is 4.70 N, he cable equivalent stiffness is 1.80N/mm and the contact profiles are the same of Prototype A. The upper segment is hinged to a horizontal pin passing through its centroidal point, while a constant weight of 1.00 N is hanged to the lower segment. Then, the upper segment is gradually rotated, so that the self-weight of the lower segment, equal to 0.08 N, together with the hanging weight, produces a moment with respect to the pitch point of the contact joint, which can be estimated by geometrically measuring the lever arm. Despite the many sources of inaccuracy (friction between profiles, friction in the sheaths, imperfections and manufacturing tolerances, creep of the tendon, parasitic bending stiffness of the tendon, measurement errors), the graph reported in Fig. 21(b) confirms the good agreement between the experimentally-measured points and the theoretical prediction, in particular for what concerns the loading\u2013unloading paths and the consequent \u2018\u2018pseudoelastic\u2019\u2019 response. The physical model of Prototype A with three bi-stable joints (\ud835\udc5b = 4 segments) is tested in the soft device under four-point-bending. Structural parameters for this case are \ud835\udc410 = 5.6N, in order to span larger values of the bending moment, and \ud835\udc3e = 1.00N/mm, because the cable is a bit longer than in the hard-device-tested beam due to the new design of the end segments" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001429_asjc.1141-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001429_asjc.1141-Figure3-1.png", "caption": "Fig. 3. The free body diagram of a 3-DOF planar mobile robot.", "texts": [ " In real applications, if the friction problem is not solved appropriately, the control accuracy will not be acceptable. In the benchmark mobile robot of this study, the effect of friction force is modeled by incorporating a coulomb and a viscous part to the total resisting force, i.e. f R t\u00f0 \u00de \u00bc Coloumb \u03bcc t\u00f0 \u00dem t\u00f0 \u00deg|fflfflfflfflfflffl{zfflfflfflfflfflffl} Coloumb \u00fe Viscous \u03bcv t\u00f0 \u00de vk k|fflfflfflffl{zfflfflfflffl} Viscous (23) where \u03bcc and \u03bcv are coulomb and viscous friction coefficients, respectively; mg is the weight force of the robot; and \u2016v\u2016 is the absolute linear velocity of the robot. In Fig. 3, the free-body diagram of the mobile robot is illustrated. This includes the control forces (Fx,Fy), control moment (\u03c4), friction resistance force ( fR), yaw angle (\u03c6), body-attached coordinate (Oxy), inertial parameters (m, J), and the inclination angle of the linear velocity (\u03d1). According to the second law of Newton and the D\u2019Alembert lemma, the equations of motion of this robot can be written as follows Fx f R t\u00f0 \u00decos \u03d1\u00f0 \u00de \u00bc m t\u00f0 \u00de _vx _\u03d5vy Fy f R t\u00f0 \u00desin \u03d1\u00f0 \u00de \u00bc m t\u00f0 \u00de _vy \u00fe _\u03d5vx \u03c4 \u00bc J t\u00f0 \u00de\u20ac\u03d5 ; \u03d1 \u00bc tg 1 vy vx 8>><>>: (24) 3" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure11.11-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure11.11-1.png", "caption": "Figure 11.11 Milling chuck for screwed-shank cutters", "texts": [ " Cutters which are used close to the spindle, such as shell end mills, are mounted on a stub arbor, Fig.\u00a011.10. This arbor is located, held, and driven in the spindle in the same way as a standard arbor. The cutter is located on a spigot or stub and is held in position by a large flanged screw. Two keys on the arbor provide the drive through key slots in the back face of the cutter. Drive key Key slot Figure 11.10 Milling-machine stub arbor Cutters having screwed shanks are mounted in a special chuck, shown in Fig.\u00a011.11. The collet, which is split along the length of its front end and has a short taper at the front, is internally threaded at its rear end. Collets of different sizes are available to suit the shank diameter of the cutter used. The collet is inserted into the locking sleeve and the assembly is screwed into the chuck body until the flange almost meets the end face of the body. The simplest method of holding a workpiece for milling is to clamp it directly to the worktable. Adequate tee slots are provided for this purpose" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001156_0954405413506587-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001156_0954405413506587-Figure3-1.png", "caption": "Figure 3. Micro-model of the surface contact.", "texts": [ " The G-W model and M-B fractal contact model were concerned only with the rough planes contact without involving rough surfaces. Specifically, these models only considered the impact of the contact surface microscopic morphology on the contact stiffness and damping without considering the impact of the macroscopic surface shape on the contact stiffness and damping. When two rough surfaces come into contact, the actual contact area is less than the one of the same size rough plane due to the impact of the surface shapes. The surface impact coefficient was introduced to account for this phenomenon (see Figure 3)34 l= ShP S Xh \u00f09\u00de where Sh is the theoretical contact area of the two surfaces which can be obtained according to different contact styles (see Table 1), P S is the sum of the areas of the two surfaces, and Xh is a synthetic curvature coefficient which is defined as Xh = 1 R1 + 1 R2 \u00f010\u00de where F is the normal load between the surfaces, E is the equivalent elastic modulus, R1 and R2 are curvature at NANYANG TECH UNIV LIBRARY on April 25, 2015pib.sagepub.comDownloaded from radius of two surfaces, respectively" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002523_j.matpr.2020.05.322-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002523_j.matpr.2020.05.322-Figure1-1.png", "caption": "Fig. 1. Modelling of Anti roll bar.", "texts": [ " To isolate the car from the uncomfortable vibrations transmitted from the road via the tyres and to transmit the manipulated forces again to the tyres so that the driver can maintain the vehicle underneath and manage to control the steady path for the vehicle. This take a look at with the roll stability, yaw balance, directional balance, body roll, experience peak, anti- bar roll. Essentially you are growing the overall spring fee or price of your suspension in the course of cornering [1]. More charge means much less body roll. Less frame roll approach much less time and a greater degree of tire contact. Fig. 1 shows the Modelling of Anti roll bar. The foremost gain is the weight reduction that may be completed through the use of hole technology. For the suspension designer, the ability of the usage of a tube substantially increases the capacity to clearly \u2018quality-music\u2019 the layout. All one has to do is change the wall thickness of the fabric to provide an incremental alternate in stiffness. A clothier can then adjust the stiffness of a bar using the equal car packaging in this manner for numerous one of a kind bars" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002942_0954406220974058-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002942_0954406220974058-Figure1-1.png", "caption": "Figure 1. Weld joint design (a) square butt joint (b) V groove edge joint.", "texts": [ " Therefore, expectation and control of welding distortion has basic significance.24 Hence, more examinations of distortion and residual stresses are required to comprehend and appreciate the advantages of A-TIG welding over TIG welding.25 The goal of the current investigation is to quantify and look at residual stresses, distortion, and thermogravimetric analysis in Inconel 625 weldment. The material used was an Inconel 625, 6.5mm thick plate. The chemical composition of Inconel 625 is given in Table 1. Figure 1(a) mentions a single-sided weld joint design with square butt edge preparation in single pass with flux coating. Multicomponent oxide flux (a combination of SiO2, TiO2, V2O5, Co3O4, and MoO3) was made into a glue by blending flux powder with acetone. Activated flux of 15-20 mm thickness was applied on the Inconel 625 plate surface before welding. Figure 1(b) mentions a weld joint design with V groove edge preparation used for TIG welding in multi pass with an ERNiCrMo-3 filler rod. Two Inconel 625 specimens of 200 200 6.5 mm3 size were readied for both welding processes. Then A-TIG and TIG weldings were completed by utilizing the parameters given in Tables 2 and 3. The chemical composition of the filler rod is presented in Table 4. The micrograph of the as-received base material (Inconel 625) is shown in Figure 2. The tensile test sample of the as-received base material was prepared and tests were carried out at room temperature as per ASTM: E8/8M standards using UTM (50 kN load cell) with the nominal loading rate kept constant at 1mm/min" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002475_s40430-020-02295-5-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002475_s40430-020-02295-5-Figure2-1.png", "caption": "Fig. 2 Schematic diagram of field point coordinates", "texts": [ " Generally, sound pressure level (SPL) is the main index used to assess and analyze sound radiation. The SPL is defined as where SPL(P) is the SPL of the field point P and pref is the reference sound pressure. In order to analyze the sound field distribution characteristics of FCACBBs, an arbitrary field point is selected to calculate its SPL. At a certain preload, the centers of all ceramic balls are assumed to be in the same plane, which is defined as the bearing plane. The bearing axis is perpendicular to the bearing plane. As shown in Fig.\u00a02, the bearing plane is placed vertically. Assuming that the bearing outer ring is fixed and the bearing inner ring rotates with the speed of n, a cylindrical coordinate system is applied to analyze the SPL of the sound field. The origin of the coordinate system lies at the intersection point between the bearing axis and the bearing plane. The 12 o\u2019clock direction is defined as position angle with 0\u00b0. The bearing rotation direction is the positive direction of the position angle. The location of field point P can be expressed as (8)p(P) = e \u22c5 + e \u22c5 n (9)SPL(P) = 20log p(P) pref Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:311 1 3 311 Page 6 of 16 where \u03b8 represents the position angle of the field point, r represents radial distance from the field point to the bearing axis, and l represents the axial distance from the field point to the bearing plane", "\u00a018, the curve of DL fluctuates in the axial distance from 20\u00a0mm to 200\u00a0mm, and the change is not obvious, which indicates that the change of the directivity can be ignored. However, as the axial distance increases, the directivity of the sound field still tends to weaken. In the above section, we analyzed the distribution characteristics of radiation noise of a FCACBB in the circumferential, radial, and axial directions, respectively. In this section, the characteristics of the whole sound field of bearings radiation noise are discussed as a whole. According to Fig.\u00a02, the cylindrical coordinate system is transformed into the rectangular coordinate system. The origin of rectangular coordinate system is consistent with that of cylindrical coordinate system. The x-axis coincides with the zero degree position angle and the positive direction is upward, while the z-axis coincides with the bearing axis and the positive direction is outward. The y-axis is perpendicular to the other two coordinate directions, and the coordinate system conforms to the right-hand rule of the Cartesian coordinate system" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001195_cp.2014.0903-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001195_cp.2014.0903-Figure1-1.png", "caption": "Figure 1.Three-stage WT gearbox structure.", "texts": [ " KF\u03b2 is the face load factor and KF\u03b1 is the transverse load factor. The factors are defined in reference [6]. Ti is the input torque of sun or each pinion being considered. A case study of a 2WM wind turbine (WT) with a gearbox of three stages, which include low-speed shaft planetary stage, intermediate-speed shaft parallel stage and high-speed shaft parallel stage, is carried out in this paper. The gearbox is assumed to have no mechanical loss. The turbine data are taken from GH Bladed [7] 2WM turbine, presented in Table 1. Figure 1 illustrates the gearbox structure being considered in this paper. Table 2 and 3 present the gearbox layout and geometry parameters for each gear stage. 2.3Fatigue Analysis Fatigue is a general phenomenon occurring in most materials with the loss of strength and other important mechanical properties as a result of cyclic loading over a period of time. Usually, components are not destroyed by a single large load, but by the accumulation of many small loads. SN -Curve is usually used to illustrate the fatigue of a material [9]", " Figure 4 shows rainflow counting result of the bending stress cycles for pinion at only intermediate stage under different mean wind speeds (8-10m/s as shown in the figure) with turbulence scale of B (a)-(c) and C (d)-(f). With S-N curve adopted from [9] and from figure 3 (under 9m/s with A class turbulence), no damage is caused in all pinions at gearbox three stages. However, according to further results of 10m/s with A class turbulence, damage is obtained for pinions at intermediate and high speed stage. The result indicates the impacts of gearbox geometry to its fatigue. In addition, from the figure 1 (e) and (f) it is observed that the aerodynamic torques under 8m/s and 9m/s mean wind speed are both in the partial load area, while under 10m/s it has a wider probability density spans area of partial load and full load. This fact indicates the reason for the damage fatigue. However, at mean wind speed above 13m/s the aerodynamic torques reversely have narrow probability density spans as blow 8m/s situation, which indicate smaller damage fatigue. From figure 4, which shows the rainflow matrix of the pinion at the intermediate-speed stage with increasing turbulence scale, and based on miner\u2019s rule, fatigue life of pinions are quantified" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003483_j.jmapro.2021.03.054-Figure17-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003483_j.jmapro.2021.03.054-Figure17-1.png", "caption": "Fig. 17. Heat transfer rates and grain sizes due to different built-up directions of the cylindrical specimens for tensile tests. The built-up direction was perpendicular to the paper for figures (a) and (d). (a) The specimen fabricated by keeping the axis along the X or Y directions; (b) Side view of Figure (a); (c) Scale showing the heat transfer rates and prior \u03b2 grain sizes; (d) The specimen fabricated by keeping the axis along the Z direction.", "texts": [ " Since the martensitic transformation variations are not timedependent, the cooling rate differences between the core and the surface in the SLM process had no significant influence on the microstructural differences. On the other hand, the EBM process went through the annealing effects, so that the cooling time influenced the microstructural properties and thus they changed from core to surface. This variation of the cooling rate depends on the size of the scanning surface. As can be seen from the diagram in Fig. 17, the X- and Y-axis samples I. Drstvens\u030cek et al. Journal of Manufacturing Processes 65 (2021) 382\u2013396 had a larger scanning area (Fig. 17(a)) than the Z-axis sample (Fig. 17 (d)). The difference in area is significant as their dimensions are small considering the scanning strategy. Only the near-surface area of the Xand Y-axis samples was affected by the slower cooling rate, and most of the core transferred the heat in the axial direction. A slower cooling rate occurred in the tensile specimens that had a cylindrical axis in the Z direction. This was mainly due to the heat flow downwards because of the small diameter. Therefore, most zones of the Z-axis specimens were affected by the slower cooling rate. A slightly faster cooling rate occurred only in the middle zone, because this part can transfer the heat both downwards and radially outwards. As mentioned above, the slower cooling rate (as it occurred near the surface) triggered CDRX in the EBM process, while a slightly higher cooling rate in the core area prevented the growth of \u03b2 grains. Therefore, the growth of \u03b2 grains, as shown in Fig. 17(c), does not show the usual phase transformation reaction. The slower cooling rate in the thermal mechanism of EBM results in a smaller grain size due to the CDRX effect. Therefore, smaller grains can be observed in the surface area of the samples on the X- and Y-axis, while in the samples on the Z- axis, smaller grains are likely to be observed in most locations except near the central area (Fig. 17d). As the tensile samples were machined (turning) from the cylindrical samples, the surface areas with finer grains were cut out in the X- and Y-axis samples, while the Z-axis samples still contained the smaller and medium grains. The microstructural results show that the core area contained basketwoven \u03b1 - laths and only a few \u03b1\u2019 martensite. Therefore, most of the X and Y axis samples contained this structure, which strengthened them. Since the grains are larger here, the slips of the lath edges were longer, therefore the elongations were lower" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002382_012046-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002382_012046-Figure1-1.png", "caption": "Figure 1. KUKA KR500 robot kinematic model.", "texts": [ " Section 3 proposes a regular sampling point planning method based on space grid, and a variable parameter identification method of kinematics and stiffness parameters is put forward. A model-based positional error compensation method is proposed in section 4. The results of the experimental verification are shown in Section 5. Finally, the paper is concluded in Section 6. Kinematic model is the foundation of robot stiffness identification and modeling. The link frames and theoretical kinematics model of KUKA KR500 industrial robot can be established through DenavitHartenberg (D-H) model method, as shown in figure 1 [11]. According to the robot kinematic model, the transformation relationship between the robot positional error and the parameters error of each link can be obtained; conversely, the parameters error of each link of the robot can be solved iteratively according to the positional error, and the actual kinematic parameters of the robot can be calibrated. Define the position error of i-th target point without external load as \u2206pi: T ,i i i i i i p p p p p a d J X X a d a d (1) CCEAI 2020 IOP Conf" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002204_0954406219900219-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002204_0954406219900219-Figure1-1.png", "caption": "Figure 1. Finger seal and detailed components: (a) laminates and cross section, (b) three-dimensional exploded view and (c) staggered laminates.", "texts": [ " First, a frictional heating model is established to analyses the radial stiffness and frictional heating. Second, a porous media fluid dynamics and heat transfer model is proposed to achieve a numerical simulation of the leakage and heat transfer characteristics of contacting FSs. Finally, the thermal deformation state is numerically analysed by using the finite element method. The contacting FS presented in this paper is composed of several finger laminates sandwiched between the spacer and forward and aft cover plates, as shown in Figure 1(a) and (b). Each finger laminate has been machined to create a series of slender flexible curved finger beams around its inner diameter, and the free end of the finger beam is called the finger foot. The finger laminates are staggered and indexed in such a way that the fingers of one laminate cover the interstices between the fingers on the adjacent laminate, which blocks axial leakage, as shown in Figure 1(b) and (c). In addition, the flexible fingers can bend radially to accommodate the rotor excursions and the relative growth of the seal and rotor that result from rotational forces and thermal mismatch without damaging the integrity of the seal, which also gives the FS a longer life than conventional labyrinth seals. The FS structure used in this paper contains fourth finger laminates with arc-moulded beams. The main structure parameters and their values are shown in Figure 2 and Table 1, respectively, where Db is the diameter of the finger base, Df is the diameter of the upper finger foot, Di is the inner diameter of the FS, Dcc is the diameter of the finger beam arcs\u2019 centres, Rs is the arc radius of the finger beam, Is is the width of the gap between fingers, Ib is the width of the finger, Lst is the length of the finger, is the finger repeat angle, 0 is the finger foot repeat angle and b is the thickness of the laminate" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002929_s12206-020-1014-6-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002929_s12206-020-1014-6-Figure5-1.png", "caption": "Fig. 5. Friction force direction and their corresponding moment arm of helical gear with geometric eccentricity.", "texts": [ " However, in the current study, the average of time-varying friction coefficient is adopted [8]. The contact pressure is simplified as the uniform distribution along the contact line. Thus, the friction force is evenly distributed along the contact line, which can be regarded as the concentrated force of the action point at the midpoint of the contact line of each segment. Compared with the simulation value, the actual friction force is smaller in the tooth tip and the tooth root but larger near the pitch point, where the friction force is zero. Fig. 5 presents the schematic of the frictional direction and the frictional force arm. Points O1 and O2 are the rotational centers, points O1\u00b4 and O2\u00b4 are the geometric centers, and the coordinate system OXY is a fixed coordinate. The normal forces on the contact line of each segment can be expressed as ( ) =ij ij FF L L t , (20) where Lij is the instantaneous length of the ith contact line and the jth segment. In Fig. 5, r1 and r2 are the instantaneous pitch radii of driven and driving gear, respectively. 1 2 1, .Pr x r l r= = \u2212 (21) The distance from point C to point P (LCP) and the distance from point P to point D (LPD) can be expressed as 2 2 01 01 AC( ) . CP P PD CD CP L x x y L L L L = \u2212 + \u2212 = \u2212 (22) The friction force of the contact line in each segment can be expressed as 1 2( ) ( ) = \u2212fi i i FF L L L t \u03bc (23) where \u03bc is the average coefficient of friction on surfaces on gear teeth. Li1 and Li2 are shown as follows: 1 2 1 2 1 1 2 1 2 0 0 ( ) / sin tan tan tan 0 0 tan . 0 i i CP i i i i CP b CP i CP b i i i i i CP b i CD b i i i CD b i L D L L L L D L L D L b L L L L L L b D L b L L D L b L \u03b2 \u03b2 \u03b2 \u03b2 \u03b2 =\u23a7 \u2264 <\u23a8 =\u23a9 = \u2212\u23a7 \u2264 < +\u23a8 = \u2212\u23a9 =\u23a7 + \u2264 < +\u23a8 =\u23a9 =\u23a7 \u2265 +\u23a8 =\u23a9 (24) The total friction is shown as 0= =\u2211 n f fi i F F . (25) In Fig. 5, S is the projection of the line O1P in the direction of the meshing line, and it can be expressed as 1 sin=S r \u03c8 . (26) The friction torque of the contact line in each section can be expressed as 2 2 1 1= \u2212fi fi i fi iT F S F S , (27) where Sij is the frictional arm, and it can be expressed as 1 2 2 1 1 2 2 0 0sin 2 sin 2 tan sin 2 i i CPi b i CP i b i CP i CP b i b i S D LLS S L LS S L D L b LS S \u03b2 \u03b2 \u03b2 \u03b2 =\u23a7 \u23aa \u2264 <\u23a8 = \u2212 +\u23aa\u23a9 \u23a7 = +\u23aa\u23aa \u2264 < +\u23a8 \u23aa = \u2212 \u23aa\u23a9 (28) 1 1 2 1 2 sin tan tan2 0 0 tan . 0 i b i PD CP b i CD b i i i CD b i LS S L L b D L b S S D L b S \u03b2 \u03b2 \u03b2 \u03b2 \u23a7 = + \u2212\u23aa + \u2264 < +\u23a8 \u23aa =\u23a9 =\u23a7 \u2265 +\u23a8 =\u23a9 The total friction torque is given as follows: 0= =\u2211 n f fi i T T " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000871_12.2185196-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000871_12.2185196-Figure2-1.png", "caption": "Figure 2. Basic coordinate system for a quadrotor.", "texts": [ " Only the most significant forces acting on the system are considered, which include: gravity, the thrusts generated by the rotors, and the reaction moments of the rotors due to aerodynamic drag. Consider a body fixed frame ( , , ) and a world inertial frame ( , , ). The location of the body frame within the world frame is given by a position vector and its orientation with respect to the world frame defined by a set of three Euler angles ( , , ), representing the pitch, roll, and yaw, respectively (as per Figure 2). Consider a ZYX rotation matrix to express the attitude of the body frame with respect to the inertial frame: (1) where c and s refer to cosine and sine, respectively. Proc. of SPIE Vol. 9468 94680R-3 Downloaded From: http://proceedings.spiedigitallibrary.org/ on 07/20/2015 Terms of Use: http://spiedl.org/terms Consider the basic equations of motion by balancing forces and moments. Assume that the center of gravity of the body is positioned in the same plane as the quadrotor arms. This assumption simplifies the equations, although it is not without its drawbacks, as discussed later" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001847_jctn.2016.4894-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001847_jctn.2016.4894-Figure3-1.png", "caption": "Fig. 3. 6 degrees of freedom arm frames.", "texts": [ " Begin Put group of ants of m; ant k goes selecting the next point of the grid; when fit Fk < fit(Goal) for each ant k do selecting the next point; if ant k arrives the destination then put a new ant in m; break; end; update pheromone trails; end; the maximum number of this group of ants is arrived; End 6. ROBOTIC ARM CONTROL EXPERIMENT 6.1. Establishment of Experimental Manipulator Avoiding Obstacle Path-Planning Objective Function with Screw Theory A 6 DOF physical mechanical arm as Figure 2 shows. Sketch frames are established in Figure 3. 6.2. Set Position and Orientation of the End-Effector Establish the base frame in Figure 7. Under the base frame and working environment of the robot, set obstacle 1, Fig. 5. Fuzzy ant colony algorithm optimization in trends. J. Comput. Theor. Nanosci. 13, 922\u2013927, 2016 925 Delivered by Ingenta to: Chinese University of Hong Kong IP: 95.181.183.70 On: Wed, 22 Jun 2016 09:25:52 Copyright: American Scientific Publishers R E S E A R C H A R T IC L E Robot motors controller PC computerManipulator Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000566_ab5b65-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000566_ab5b65-Figure15-1.png", "caption": "Figure 15. The spalled tooth on the ring gear.", "texts": [ " By embedding the local periodicity, PRPCA presents its superior ability on extracting complete fault information, which can be seen in figure\u00a014(d). The three pairs of impulse spaced by 13 teeth coincide with the defects on the gear. This provides the evidence that PRPCA is prone to periodic anomaly only. Ring gear is usually an integrated part in the gearbox casing and it is non-replaceable during the whole life cycle of the gearbox. In this subsection, an experiment in terms of spalled tooth failure on the ring gear is invested. The damaged ring gear is exhibited in figure\u00a015. After the difference operation, some abnormal components can be observed from the IAA signal, which is shown in figure\u00a016(a). However, we can hardly conclude that the ring gear is broken according to these exceptions because of the existence of environmental factors, such as the fluctuation of load and external knocks. Figure\u00a016(b) displays the result obtained by MTSA-PRPCA, there is no doubt that the anomaly induced by fault is detected successfully. Since there are three planet gears in the gearbox, the spalled tooth is meshed three times within a cycle and each generates an impulse, which validates the correctness of the result" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002475_s40430-020-02295-5-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002475_s40430-020-02295-5-Figure5-1.png", "caption": "Fig. 5 Experimental scheme", "texts": [ " The cooling temperature was set at a constant of 18\u00a0\u00b0C. The engine oil with 30# was used for bearing lubrication. The pressure of air was 0.28\u00a0MPa. The flow rates of lubrication oil and cooling water were set to 0.025\u00a0mL\u00a0min\u22121 Journal of the Brazilian Society of Mechanical Sciences and Engineering (2020) 42:311 1 3 311 Page 8 of 16 and 5.0\u00a0L\u00a0min\u22121. The preload of bearings was adjusted to 350\u00a0N. The motorized spindle was driven by built-in motor, and it operated at a steady state during the test. As shown in Fig.\u00a05, the 24 measuring points were selected to verify the accuracy of the calculation and analyze the radiation noise of the FCABB applied in the motorized spindle. The axis of the spindle (bearing) was placed in the horizontal plane. All points were arranged in a plane of the front of the spindle (bearing), and the plane was 50\u00a0mm away from the bearing plane. The measuring points were evenly distributed in the circumferential direction with 210\u00a0mm away from the center line of the bearing. The angle interval between the adjacent two measuring points was 15\u00b0" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002866_0954406220971666-Figure6-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002866_0954406220971666-Figure6-1.png", "caption": "Figure 6. Operating modes of the stator in the simulation: (a) Mode A. (b) Mode B.", "texts": [ " Moreover, the material of the piezoelectric elements is PZT-8H, whose piezoelectric constant matrix and elastic stiffness matrix are also demonstrated in Table 2. In addition, the fine thread in this configuration makes it difficult to complete the modal analysis of the stator. Since the influence of internal threads on the results of modal analysis is so small that can be ignored, the computing model thus is simplified as a circular cylinder to stimulate under the free boundary condition. The vibration modes of the stator by FEM are shown in the Figure 6. The resonant frequency of mode A and B are 12.584 kHz and 12.613 kHz respectively. The frequency difference between two orthogonal modes is 29Hz, which can substantially meet the design requirement of the motor. Dynamic response analysis. To validate the dynamic features of the motor, the dynamic response analysis is constructed in a damping coefficient of 0.3%. Based on the modal analysis results, the exciting signals with the peak voltages of 80V under the frequency of 12.584 kHz, are applied on the piezoelectric elements as illustrated before" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003508_j.optlastec.2021.107126-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003508_j.optlastec.2021.107126-Figure7-1.png", "caption": "Fig. 7. 3D plot - Interaction effects of process variables on case depth.", "texts": [ " 6(a) and (e) with same scanning speed and focal position, at LP 1000 W the hardened depth is G. Muthukumaran and P.D. Babu Optics and Laser Technology 141 (2021) 107126 2.3 times higher than that of LP 500 W. Maximum case depth of 1.39 mm achieved at LP 1000 W, SS 240 mm/min and D 295 mm, shown in Fig. 6 (f). Along with scanning speed and laser power, focal plane position also influences the case depth of the hardened region. Response surfaces of case depth in terms of process input variables is shown in Fig. 7. Case depth variations can be explained with respect to interaction time and laser beam density; Table 9 depicts the case depth variations. Comparing #Runs 1 and 6, for the same laser beam density with different interactions time 0.64sec and 0.43sec, the variations in case depths are marginal. With minimum beam density value, 16.60 W/mm2 (LP 500 W) with interaction time of 0.5sec case depth achieved is 0.53 mm. The maximum case depth achieved with LP 500 W is 0.68 mm, and this is attained with the focussed condition. From the trial, #Runs 4,8 and 12, a laser density of 39.20 W/mm2 and interaction time 0.35sec case depth achieved is 0.89 \u00b1 0.15 mm. From trial #Runs 7 and 15, it is understood that interaction time plays a dominant role in case depth formation. Fig. 7 illustrates the case depth of the hardened region based on the input process variables. Fig. 7(a) shows the relationship between LP and SS on the hardened layer\u2019s depth. The relation between heat input and the process inputs is shown in Eq. (6). Maximum heat input at the surface can be achieved with high LP, and low SS. Low SS means high interaction time, and more interaction time give rise to high case depth. Fig. 7(b) depicts the case depth behaviour with LP and FPP. Laser beam intensity varies with the FPP; from Table 9, it can be concluded that beam intensity at the focusing position on the surface yields high case depth when compared with other defocussed positions. Enlargement of beam geometry dimensions at defocused position results in low beam densities, which affects the bead geometry. Fig. 7(c) depicts the behaviour of the case depth concerning SS and FPP. From Table 8, it is evident that at the focusing position, the laser beam coverage area is minimal. It increases with an increase in FPP; this depicts that the laser beam density is very high at minimum laser beam coverage area. By defocusing the focal plane position and reducing the scanning speed, case depth will increase, and this statement is substantiated by trial #Run 15. Laser power, scanning speed and focal plane position and material properties are the major considerations in single-pass laser hardening to achieve desired surface properties" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003527_j.tafmec.2021.102991-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003527_j.tafmec.2021.102991-Figure8-1.png", "caption": "Fig. 8. Loading and boundary conditions considered in [29].", "texts": [ " But, with complicated geometry, it is difficult to estimate the accurate value of SIF through the analytical method. The SIF determined in the present study is validated with the published numerical work [29]. It is observed that the SIF is in good agreement with the experimental investigation [20]. A 3D symmetric spur gear is developed and a uniform distributed load is applied along the face N.V. Namboothiri and P. Marimuthu Theoretical and Applied Fracture Mechanics 114 (2021) 102991 width. The gear parameters are given in Table 2. The loading and boundary conditions are shown in Fig. 8 (see Fig. 9). A twenty noded quadratic brick element is adopted to discretize the fillet region and fifteen noded triangular prism element is used to discretize the crack region. The fracture analysis is performed using the contour integral method. The face is divided into 28 elements and the load F = 1783 N/mm is applied on each node along the face width. The KI, KII and KIII at each node on crack tipis determined and the results are compared with the published numerical study [29] (Fig. 10). It is observed that a deviation of 11" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000228_j.promfg.2019.04.046-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000228_j.promfg.2019.04.046-Figure1-1.png", "caption": "Figure 1. The five steps in the spectral graph theoretic approach used to estimate the heat flux in the part layer-by-layer. Here we show an embodiment of the laser powder bed fusion (LPBF) process.", "texts": [ ", the beam diameter and shape, and subsequent diffusion on the powder bed surface are not accounted. The time required by the laser to fuse a layer is infinitesimal compared to the time taken to deposit a new layer. Hence, heat dissipation occurs only during the powder deposition process as the bed is lowered and the recoater applies a new layer. These assumptions can provide a more comprehensive model, which will be pursued in later works by the authors. The approach has the following five steps, as shown in Figure 1. Step 1: Graph Embedding - Constructing a network graph from the nodes as a layer is deposited by the recoater. The part is sliced into layers and hatches, and a fixed number of spatial locations (nodes) which are randomly sampled. The number of nodes is contingent on the geometry of the part, in this work, a density of 5 nodes per mm2 provided a sufficiently good approximation to the heat flux estimated with a moving heat source solution obtained through FEA [18, 19]. Step 2: Constructing a network graph from the nodes as a layer is deposited by the recoater" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001677_s12206-015-0718-5-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001677_s12206-015-0718-5-Figure2-1.png", "caption": "Fig. 2. The contact model between the worktable and the guide way.", "texts": [ " The model is created by mode superposition, and RungeKutta method is used to analyze the dynamic response of the system, including the vibration of the screw, the axial vibration amplitude of the cutter\u2019s work point, and also the stability of the drive system. 2. Dynamic model of the system To consider the contact deformations, the bearings at both ends of the screw and the contact between the screw and the nut are expressed by six stiffness and damping coefficients respectively. Fig. 1 is the bearing\u2019s contact model, and the contact model between the screw and the nut is the same as Fig. 1, but the parameters are different. The contact model of the worktable and guide way is shown in Fig. 2. All parameters are listed in the nomenclature. In the system, the screw rotates at a constant angular veloc- ity \u03a9 . The coordinate origin is at the left end of the screw. The coordinate system and the dynamic model of the system are shown in Fig. 3. The worktable translation ( )s t in x direction includes axial translation produced by the screw rotation, axial one mlj caused by torsional deformation of the screw, ( ),du tlq from axial deformation of the screw and xq from local vibration of the worktable, and is expressed as: ( ) ( ),d m d xs t u t qlq lf lq= + + + " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000043_s12541-019-00047-7-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000043_s12541-019-00047-7-Figure8-1.png", "caption": "Fig. 8 Sketch map of the shock velocity on a the section and b the normal plane", "texts": [ " Likewise, if SRNPE is negative, teeth enter meshing earlier, which causes the transmission error curve to move up. Referring to this literature, we can obtain the unloaded transmission error considering SRNPE, which is the red curve shown in Fig.\u00a07. In this figure, the blue curve is the loaded transmission error considering SRNPE, and the black circles are the actual meshing-in positions. Because the actual contact position deviates from the theoretical line of action, the linear velocities between the pinion and the gear are unequal, and the difference is the impact speed. Figure\u00a08 shows the calculation principle of the impact speed. M is the actual meshing-in position; M1M2 is the instantaneous line of action; N1N2 is the theoretical line of action; vp and vg are the circular speeds of the pinion and the gear, which are projected into M1M2 respectively, noted as vnp and vng; rbp and rbg are the theoretical base radii; and rbg\u02b9 is the instantaneous base radius. The impact speed vs can be computed by Eq.\u00a0(4). where vsi is the impact speed; RMi and nMi are the position vector and normal vector of M; \u03c1Mi is the vector from Og to M, with \u03c1Mi= RMi-E; E is the vector from Of to Og; wp and wg are the angular speeds of the pinion and the gear" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002258_s00170-020-04953-3-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002258_s00170-020-04953-3-Figure4-1.png", "caption": "Fig. 4 Virtual rotation center grinding principle", "texts": [ " The generating plane rotating around the gear axis O2, \u03c92 means the theoretical angular velocity, when line segments O2M, O2N and MN are represented by vectors rM, rN, and rNM, respectively. The velocity can be expressed as: vN \u00bc vM \u00fe \u03c9T rNM \u00f08\u00de where vN = \u03c92 \u00d7 rN and vM = \u03c9M \u00d7 rM, while \u03c92 \u00d7 rN =\u03c9M \u00d7 rM +\u03c9T \u00d7 rNM. Here, \u03c9M is the equivalent angular velocity of the pointM. And the following conclusions can be determined: \u03c92 \u00bc \u03c9M \u00bc \u03c9T \u00f09\u00de The tooth surface of planar enveloping hourglass worm is ground by the plane grinding wheel, as shown in Fig. 4. The grinding wheel plane is tangent to the main basic circle and point o2 is actual rotation center. In order to grind the planar enveloping hourglass worm tooth surface, the hourglass worm should be rotated around the C axis with angular velocity \u03c9c, and the plane grinding wheel should be rotated around point o2 with angular velocity \u03c92, and the angular velocity \u03c9c and \u03c92 need to meet the following condition: \u03c9C \u03c92 \u00bc i12 \u00f010\u00de Based on the aforementioned definition, the virtual rotation center theory is proposed as shown in Fig. 4. Here, the motion of rotating around actual rotation center is decomposed into the motion of X axis, Z axis, and B axis linkage. The velocity vectors vX and vZ indicates the rectilinear motion velocity vector of the workbench on the X axis and Z axis, respectively. The angular velocity vector \u03c9B shows the rotational motion angular velocity vector of the workbench on the B axis. In order to assure the grinding wheel plane is always tangent to the main basic circle, the velocity vector and angular velocity vector should satisfy the following conditions: \u03c9B \u00bc \u03c92 vX \u00fe vZ \u00bc \u03c92 ro2B \u00f011\u00de According to the Eqs. (10) and (11), the kinematics relation of four motion axes can be determined. In the grinding process, the B axis and point o3 is actual rotation center of the workbench, while point o2 is virtual rotation center of the workbench. The actual grinding center distance is a' and the theory center distance expresses by a. In Fig. 4, it can be known that the grinding center distance a' is much smaller than the theory center distance a. Therefore, the machine structure with this virtual rotation center grinding principle is much smaller than the actual rotation center grinding principle. Based on the virtual rotation center grinding principle the four-linkage grinding movements are shown in Fig. 5. Here, point P1 is the actual rotation center of the workbench at the beginning of grinding, while point P2 represents the actual rotation center of the workbench at the ending of grinding" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000465_0954406219878755-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000465_0954406219878755-Figure3-1.png", "caption": "Figure 3. (a) Schematic for machining enveloping conical worm; (b) coordinate systems for machining enveloping conical worm.", "texts": [ " Consequently, the equation for the grinding wheel generating surface, g, can be determined in a by the aid of the sphere vector function,4,5 and the outcome is ~ra a \u00bc u ~ma , g \u00bc u sin g cos ~ia \u00fe u sin g sin ~ja \u00fe u cos g ~ka, u4 0, 04 5 2 \u00f01\u00de in which the symbols u and are the two surface parameters of g. The unit normal vector to g can be obtained as ~n a \u00bc @ ~ra\u00f0 \u00dea @ @ ~ra\u00f0 \u00dea @u @ ~ra\u00f0 \u00dea @ @ ~ra\u00f0 \u00dea @u \u00bc ~na , g \u00bc cos g cos ~ia \u00fe cos g sin ~ja sin g ~ka \u00f02\u00de Obviously, the orientation of ~n a achieved above is from the grinding wheel entity to the space as displayed in Figure 2. Generation of enveloping conical worm As described in Figure 3, a fixed coordinate system o1 O1; ~io1, ~jo1, ~ko1 is affiliated with the worm rough. The origin O1 of the coordinate system o1 locates at the middle point of the thread length of the worm and the unit vector ~ko1 lies along the worm rough axis. A rotating coordinate system 1 O1; ~i1, ~j1, ~k1 is rigidly linked to the worm rough and is utilized to denote its present location. The rotating angle of 1 relative to the coordinate system o1 is the angle \u2019. The position of the grinding-wheel-rack is determined by the two moving coordinate systems, od Od; ~iod, ~jod, ~kod and d Od; ~id, ~jd, ~kd " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001985_1.4034768-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001985_1.4034768-Figure2-1.png", "caption": "Fig. 2 Three-dimensional model of (a) defective rod-fastened rotor and (b) defective disk b", "texts": [ " In this paper, the impacts of the flatness error of axial assembly interfaces, circumferential position error of rod-holes, and concentricity error of disks are investigated. 2.1 Static Model of 3D Defective Rod-Fastened Rotor With Flatness Error. The flatness error of axial assembly interfaces is represented by the machining defect in disk b. This error is assumed to be formed by overcutting during the milling process. The 3D static finite element model of defective rod-fastened rotor is built as follows: (1) Two shafts, three disks, and 12 rods are built by eight node elements by finite element method [8] (Fig. 2(a)), and each node in an element has three translational degrees-offreedom (DOFs) including x, y, and z. (2) The overcutting defect in disk b is determined by the angle h and depth d, which is equal 60 deg and 10 lm, respectively (Fig. 2(b)). In order to make the defect sufficiently clear to see, the defect depth in Fig. 2(b) has been amplified by 1000 times. The way of building defect depth and angle is to remove the defect part from the whole solid disk and then it is meshed by finite elements. (3) The assembly contact interfaces are also the important components of the static model. The augmented Lagrangian contact algorithm involving friction [9] is used to calculate normal and tangential stress of interfaces. This method is also used in the software ANSYS. (4) The load in this model is pretightening force (188 KN for each rod) and centrifugal force (determined by working speed 7500 rpm)", "org/about-asme/terms-of-use The displacement vector xR is written as xR\u00bc\u00f0x1; y1; z1;\u2026; xnd ; ynd ; znd \u00deT (3) where nd is the node number of 3D defective rotor; and xi, yi, and zi (i\u00bc 1, 2,\u2026, nd) are the displacements of ith node. The reduction of system equation dimensions is performed by a component mode synthesis method [15]. The reduced rodfastened rotor is described by the following equations: M\u20acq \u00fe G _q \u00fe Kq \u00bc Q (4) where T is the transformation matrix to reduce the original system; M\u00bcTTMRT; G\u00bcTTGRT; K\u00bcTTKRT; and Q\u00bcTT(KRrb \u00femx2e\u00fe g\u00fe fR)T. Then, the reduced 3D rotor system in state space is expressed as _X \u00bc _q M 1\u00f0Q Kq G _q\u00de (5) where X \u00bc \u00f0qT; _qT\u00deT is the state parameter. For the reduced rotor system in Fig. 2(a), the number of reserved DOFs is 27, which involve 15 nonlinear DOFs and 12 linear eigenmodes. The nonlinear DOFs come from five nodes (three nodes at disks and two nodes at bearings) acted by nonlinear forces. The linear eigenmodes are composed of six elastic eigenmodes and six rigid-body modes. 3.2 Nonlinear Solution Method. The mass eccentricity and rotor bending have the T-period feature. Thus, the solutions are determined by Eq. (5) and T-period boundary condition X(t)\u00bcX(t\u00fe T). The classic equations are listed as _X \u00bc f \u00f0X;t;k\u00de X\u00f0t\u00de \u00bc X\u00f0t\u00fe T\u00de ( (6) with f \u00f0X;t;k\u00de \u00bc _q M 1\u00f0Q Kq G _q\u00de , and k as the system parameter, such as e and x" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003158_9781315768229-Figure17.18-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003158_9781315768229-Figure17.18-1.png", "caption": "Figure 17.18 Vee-bending tool", "texts": [ "14 Simple blanking tool Top plate Punch plate Punch Stripper plate Die Bolster Fixed stop Figure 17.15 Use of stops in simple blanking Fixed stop (a) (b) Sliding stop Work-material strip Strip feed Figure 17.16 Follow-on tool Pilot Workpiece against it by hand, Fig.\u00a017.15(a). The punch descends D ow nl oa de d by [ U ni ve rs ity o f C al if or ni a, S an D ie go ] at 2 1: 53 2 8 Ju ne 2 01 7 17 Presswork 17 270 Vee-bending tools consist of a die in the shape of a vee block and a wedge-shaped punch, Fig.\u00a017.18. The metal to be bent is placed on top of the die \u2013 suitably located to ensure that the bend is in the correct position \u2013 and the punch is forced into the die. To allow for springback, the punch is made at an angle less than that required of the finished article. This is determined from experience \u2013 e.g. for low-carbon steel an angle of 88\u00b0 is usually sufficient to allow the metal to spring back to 90\u00b0. Side-bending tools are more complicated than those employed in vee bending but give a more accurate bend" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002058_1.4035203-Figure15-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002058_1.4035203-Figure15-1.png", "caption": "Fig. 15 The shaking and picking points", "texts": [ "org/about-asme/terms-of-use torsional vibration, respectively; cx and cy are the damping coefficients of the bearing (excluding the influence of contact friction) in the horizontal and vertical directions, respectively; kx and ky are the equivalent stiffness of the bearing in the horizontal and vertical directions, respectively; Ffy represents the vertical component of the friction force; Mf is the resultant of torque associated with the friction forces Ffx and Ffy; xs0 and ys0 are the translation displacements of the shaft in the X and Y axes, respectively, corresponding to the location of the support; l0 is the distance between the stern support and the origin of the coordinate system in the Z axis; and Prx and Pry are the sound pressure. The partial differential equations (18a)\u2013(18c) can be solved with the method applied to the support beams. 5.1 Parameter Estimation. In order to estimate the damping ratios of the natural modes of the submerged stern, the FRFs between the driving and picking points (force-to-force) were measured. The two points are both located on the horizontal beam of the stern support, as illustrated in Fig. 15. To identify those modes associated with torsional vibration, the FRFs corresponding to different shaft speeds were measured [32], respectively. The measured FRFs, which correspond to the shaft speeds of 0 rpm and 60 rpm, respectively, are shown in Fig. 16. It can be seen that the frequencies of those modes at 86.5 Hz and 113.5 Hz change with shaft speed, but the modes at 30.5 Hz and 80 Hz are not affected by rotation. Based on the results of finite element analysis and measurement, it can be concluded that those modes at 30" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002148_ecce.2016.7854981-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002148_ecce.2016.7854981-Figure1-1.png", "caption": "Fig. 1. Cross section of 1/12 model of switched reluctance motor.", "texts": [ " In this paper, the reduction of noise and vibration in magnetically saturated region, that is a typical condition in switched reluctance motors, is investigated. The current waveform is derived to minimize variation of the radial force sum with radial force approximation. Proposed driving current is consist of DC, fundamental, second, and third harmonic components. A comparison is presented among the conventional square current waveform, the proposed, and the previous current waveforms. Finally, the experimental test is carried out to confirm the reduction of the acoustic noise by the proposed current. II. RADIAL FORCE IN SWITCHED RELUCTANCE MOTOR Fig. 1 shows a part of a cross section of the switched reluctance motor. In this study, a three-phase switched reluctance motor is investigated. The number of stator and rotor poles are 36 and 24, respectively. The neighboring stator 978-1-5090-0737-0/16/$31.00 \u00a92016 IEEE poles are assigned as phase A, B, and C. Each stator pole generates radial forces, FrA, FrB, and FrC. The radial force sum Frsum is the summation of three radial forces. Figs. 2 (a) and (b) show principles of the current waveforms and radial forces" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002762_j.euromechsol.2020.104125-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002762_j.euromechsol.2020.104125-Figure2-1.png", "caption": "Fig. 2. Sketch of the interaction of the rolling element with (a) outer race defect (Luo et al., 2019a), and (b) inner race defect, (c) Load zone in REB (d) Rolling element deformation in load zone (e) Rolling element on the verge of interaction with.", "texts": [ " Then the experimental study to correlate the peak acceleration and force in the load zone is presented in the middle part of section 2. The last part of section 2 describes the signal processing based methodology to locate entry and exit events along with methods to handle any complicacy arising in it. Defect and bearing specifications, along with the experimentation carried out is presented in section 3. Section 4 discusses the results of the proposed method, and the paper is concluded in section 5. The rolling element-spall interaction shown in Fig. 1(a) is reproduced in Fig. 2 for a better illustration of the interaction. In this study, it is assumed that the rolling element does not interact with the bottom surface of the defect. The interaction of defect and rolling element, as shown in Fig. 2 (a) and (b), can be understood as the following series of events:(p, P) normal contact between the rolling element and the raceways, (q, Q) is the start of the interaction between the leading edge of the defect and the rolling element (Fig. 2(e)), (r, R) initiation of rotation of the rolling element about the leading edge, (s, S) complete destressing of the rolling element, and, (t, T) impact between trailing edge of the defect and the rolling element. Rolling element central positions A, B, and C from Fig. 1 are shown as q, s, and t respectively in Fig. 2 (a) and A.P. Patil et al. European Journal of Mechanics / A Solids 85 (2021) 104125 (b). Time spacing between rolling element center positions are defined (Luo et al., 2019a) as the time from entry to complete destressingtedand Ted, time from start to complete destressing tsd and Tsd, time from complete destressing to impact tdi and Tdi, where tis for outer race defect and is T for outer race defect. In the work of Luo et al. (2019a), the time separation tsd(or Tsd) was found out analytically using a physics-based approach, while tdi (or Tdi) was found out from recorded signal", " The rolling element center positions during the interaction and geometric relationships are shown in Fig. 3 for outer race defect. The time separation from start to complete destressing in case of outer race defect is given in Eq. (1) (Luo et al., 2019a) tsd = \u03c8b + \u03c8d \u03c9cage (1) where \u03c9cageis cage angular speed, and the angle \u03c8b (Fig. 3(b)) can be obtained from Eq. (2)(Luo et al., 2019a) \u03c8b \u2248 tan(\u03c8b)= bo ( 0.5Dp + 0.5Db ) \u2212 (\u03b4 + 0.5cl) (2) where bo is semi-width of contact area in the tangential direction (Fig. 2 (d and e) and Fig. 3(a)), Dpis pitch diameter of the bearing,Db is rolling element diameter, clis radial clearance (Fig. 3(b)), \u03b4is total deformation of rolling element-raceway contact in the radial direction. The semiminor dimension bo of contact area between the outer race and rolling element at position \u03d5 is given as (Harris, 1996) bo = b* o [ 3F\u03d5 2 \u2211 \u03c1o ( 1 \u2212 \u03be2 I EI + 1 \u2212 \u03be2 II EII )]1 / 3 (3) where b* ois dimensionless parameter dependant on the curvature of bodies in contact, F\u03d5is load acting on the rolling element at angular position \u03d5 (Fig. 2 (c) and (d)), \u2211 \u03c1o is the sum of curvature of outer racerolling element contact, \u03be is Poisson\u2019s ratio, E is the modulus of elasticity, while the subscripts I and II indicate two bodies in contact. The total deformation \u03b4 depends on the contact deformations at the contact of rolling element contact with inner and outer raceways and is expressed in Eq. (4) \u03b4= \u03b4i + \u03b4o (4) Where \u03b4i and \u03b4oare contact deformation between rolling element contact with inner and outer race, respectively. These are computed using the Hertzian contact theory and are expressed in Eq", "3 is dedicated to locating points B and C on the measured signal. For determining the angles \u03c8d and \u03b1d analytically, computation of \u03b4, bi and bo is essential, which depend onF\u03d5and was assumed to be maximum in Ref. (Luo et al., 2019a; Moazen-ahmadi and Howard, 2016). In this study, a more realistic approach to obtaining the value of force F\u03d5 is proposed by using force distribution in static condition. When the bearing is subjected to load in the static condition, some rolling elements support the load, thus forming a load zone, as shown in Fig. 2(c). The variation of F\u03d5 in the load zone is governed by Eq. (24) F\u03d5 =Fmax ( 1 \u2212 1 \u2212 cos \u03d5 2\u03b5 )n (24) where Fmaxis the maximum load in the load zone, \u03d5is the angular position of the rolling element in the load zone, and\u03b5is dimensionless load parameter defined as \u03b5= 1 2 ( 1 \u2212 cl \u0394r ) (25) where \u0394ris the radial shift in bearing due load application and cl is radial clearance. The maximum force in the load zone for ball bearing is given as (Harris, 1996) Fmax = 4.37Fr N cos \u03b1 (26) As can be seen from Eq" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001042_j.triboint.2014.08.015-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001042_j.triboint.2014.08.015-Figure1-1.png", "caption": "Fig. 1. Schematic of the rolling contact test rig.", "texts": [ " So micrometer scale grooves are engraved on the surface of a 100Cr6 bearing steel ring by two methods, laser engraving and wire cutting. A multigrid method, using a combination of methods described by [19,20], is used to predict the film thickness and pressures generated in the lubricant, for a transverse groove of varying dimensions. These numerical method results are used to explain the experimental observations of the tests which study the effect of varying wavelength on the interfacial torque. The test rig consists primarily of two drive shafts and a hydraulic cylinder Fig. 1. One test ring is mounted on each shaft respectively. The two ring surfaces are brought into contact against each other by applying a normal load, using a hydraulic cylinder. Each ring has a conical bore, fitting on to a corresponding conical shape on the shaft. Contents lists available at ScienceDirect journal homepage: www.elsevier.com/locate/triboint Tribology International http://dx.doi.org/10.1016/j.triboint.2014.08.015 0301-679X/& 2014 Elsevier Ltd. All rights reserved. n Corresponding author" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001023_robio.2014.7090430-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001023_robio.2014.7090430-Figure4-1.png", "caption": "Fig. 4. 3D model of Stewart platform.", "texts": [ " The following section explains the PHARAD mechanism and a measuring method for the ankle joint motion by solving the inverse kinematics of the parallel link mechanism. In Section 3, the attachment method for a foot and leg and the driving principle of the PHARAD are presented. Then, the efficacy of the developed device is shown by presenting experimental results in Section 4, and the conclusions are summarized in the last section. This section describes the PHARAD mechanism. We adopted Stewart platform, which is one of the parallel link mechanisms [19]. The moving plate and base plate are connected by six linear actuators, as shown in Fig. 4. The posture of the moving plate can be measured and controlled with six DOFs by measuring and changing the length of each linear actuator. By attaching a patient\u2019s foot to the moving plate, this mechanism can conduct ROM measurement and rehabilitation adapted to complex ankle motions. Fig. 5 shows the geometric model of the Stewart platform. and are the coordinate systems of the base plate (the upper plate) and the moving plate (the lower plate), and and are the origins of and , respectively. In , the position of is " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003527_j.tafmec.2021.102991-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003527_j.tafmec.2021.102991-Figure3-1.png", "caption": "Fig. 3. The CAD model of asymmetric spur gear.", "texts": [ " The gear parameters such as module, number of teeth, gear ratio and addendum factor were varied to perform the parametric study. where ra = Addendum diameter rf = root circle diameter rbd = base circle diameter at drive side rbc = Base circle diameter at coast side N.V. Namboothiri and P. Marimuthu Theoretical and Applied Fracture Mechanics 114 (2021) 102991 rHPSTC = radius of highest point of single tooth contact The gear surface is partitioned into a single tooth and double tooth contact regions (Fig. 3). The load is shared with the adjacent teeth when the gears are in contact between A to B. But the total load is acting on a single tooth when it comes in contact between B and C. Hence, this region is considered the critical region. Further, the load is shared with the adjacent teeth when the gear comes in contact between C and D. A = Highest point of tooth contact (HPTC) B = Highest point of single tooth contact (HPSTC) C = Lowest point of single tooth contact (LPSTC) D = Lowest point of tooth contact (LPTC) Ht = Tooth height Rt = Rim thickness The root crack in the spur gear initiates at the most critical region where the maximum principal stress occurs" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000553_rnc.4809-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000553_rnc.4809-Figure9-1.png", "caption": "FIGURE 9 Error trajectories when \ud835\udefc2 = 0.6 [Colour figure can be viewed at wileyonlinelibrary.com]", "texts": [], "surrounding_texts": [ "In this section, the effectiveness of the proposed method is verified by a numerical simulation. In the following, the considered MASs contain four agents indexed by numbers 1, 2, 3, and 4, their communication topology is illustrated in Figure 2. As shown in Figure 1, because the communication topology is a connected graph, Assumption 1 is guaranteed. The Laplacian matrix is obtained as \u2112 = \u239b\u239c\u239c\u239c\u239d 2 \u22121 0 \u22121 \u22121 2 \u22121 0 0 \u22121 2 \u22121 \u22121 0 \u22121 2 \u239e\u239f\u239f\u239f\u23a0; its eigenvalues are \ud835\udf062(\u2112 ) = 2, \ud835\udf063(\u2112 ) = 2, and \ud835\udf064(\u2112 ) = 4. The motion of each agent satisfies (1) with A = ( 0 1 \u22121 0 ) and B = ( 1 1 ). Because ( 1i \u22121i 1 1 )\u22121 A ( 1i \u22121i 1 1 ) =( \u22121i 0 0 1i ) , it implies that Assumption 2 holds and the matrix A has a pair of conjugate complex eigenvalues as \u00b11i. According to Theorem 1, the maximal allowable time-delay \ud835\udf0f\u2217 is allowed to be \ud835\udf0b 2 . Given matrix A and communication topology \u2112 , the maximal time-delay \ud835\udf0f\u2217 and the gain scalar \ud835\udf07 can be conveniently obtained by Theorem 1. Furthermore, given any positive real numbers \ud835\udefc1, \ud835\udefc2 and \ud835\udefd, the form of the event-trigger can be accordingly determined. By the event-trigger (7) and the consensus protocol (8), Theorem 1 says that the bounded consensus of the MASs can be guaranteed. For validating the effectiveness of the proposed method, by Theorem 1, it gives \ud835\udf07 = \u22120.05. The data sampling period is T = 0.2; thus, \ud835\udf0f\u2217 = 7T. The thresholds of the event-trigger are chosen as \ud835\udefc1 = 1.0, \ud835\udefc2 = 0.15, and \ud835\udefd = 0.5. The state initial values are supposed to be x1(0) = (1, 1)T, x2(0) = (1,\u22121)T, x3(0) = (\u22121, 1)T, and x4(0) = (\u22121,\u22121)T. Without the event-trigger (7) and the consensus controller (8), the state trajectories of the four agents are shown in Figure 2. As shown in Figure 2, the trajectories of the four agents are respectively illustrated in three dimensional coordinates. One can intuitively see form Figure 2 that the trajectories of four agents always fluctuate and the agents 2, 3, and 4 are not converged with the agent 1. It indicates that the desired consensus cannot be achieved once the control law is not employed. By employing the proposed event-trigger (7) and consensus protocol (8), the trajectories of agents xi(t) and the errors ej(t) are shown in Figures 3 and 4, respectively. FIGURE 3 State trajectories of four agents [Colour figure can be viewed at wileyonlinelibrary.com] FIGURE 4 Error trajectories of agents [Colour figure can be viewed at wileyonlinelibrary.com] Figure 3 depicts the trajectory snapshots of the four agents under the proposed event-triggered consensus protocol. One can intuitively observe from Figure 3 that the curves gradually converge together; the consensus is achieved asymptotically under the proposed scheme. As shown in Figure 4, the curves show the errors between agents 2, 3, and 4 with agent 1. It can be seen that the errors are gradually reduced down and converge within a small region. The event-triggered time instants of four agents ki, i = 1, 2, 3, 4 are, respectively, shown in Figures 5 to 8. The upper graphs in Figures 5 to 8 show the trajectories of function fi(t) and the threshold line \ud835\udefc1e\u2212\ud835\udefdt + \ud835\udefc2; the sawtooth-like lines show the trajectory of fi(t), while the smooth line depicts the trajectory of \ud835\udefc1e\u2212\ud835\udefdt + \ud835\udefc2. The value \u201c1\u201d in the lower graphs in Figures 5 to 8 mean that the event-triggers are violated at the current time instant. Every time when ||fi(t)|| is greater than \ud835\udefc1e\u2212\ud835\udefdt + \ud835\udefc2, the event-trigger i is violated and the function ||fi(t)|| is reset to zero. Afterward, the function ||fi(t)|| is gradually increased from zero until the next triggering time. One can directly observe that the four event-triggers are only violated by 14, 7, 8, and 7 times. It implies that the event-triggered frequencies of four agents are, respectively, 2.8%, 1.4%, 1.6%, and 1.4%, and over 97% of data is not broadcasted. It gives the proof that the desired bounded consensus of MASs can be achieved; meanwhile, the data transmission amount can be significantly reduced. As discussed in Remark 3, the parameter \ud835\udefc2 plays an important role in improving consensus of MASs. To further evaluate its influence, by setting \ud835\udefc2 = 0.6, the error trajectories and the event-triggered time instant of agent 1 are, respectively, shown in Figures 9 and 10. Comparing with Figure 4, it can be found from Figure 6 that the error trajectories fluctuate within a larger region. It indicates that the control performance will be deteriorated if the threshold \ud835\udefc2 is amplified. However, comparing with the former case, as shown in Figure 5, it can be seen from Figure 10 that event-triggered amount of agent 1 is reduced from 11 to 5. Since less data is broadcasted by event-trigger 1 to its neighbors, the data transmission burden is greatly alleviated. Choosing \ud835\udefc2 = 0.05, the error trajectories and the event-triggered time instant of agent 1 are, respectively, shown in Figures 11 and 12. Figure 11 shows that the error trajectories are reduced down close to zero with a quicker speed and converged within a smaller region. However, Figure 12 shows that triggered frequency of event-trigger 1 is also significantly increased. Therefore, given a smaller threshold \ud835\udefc2, the consensus performance can be improved; however, the data broadcast amount of the event-trigger will be accordingly increased. By the aforementioned numerical validation, it can be concluded that the threshold \ud835\udefc2 plays a key role in achieving the consensus of MASs. It provides a trade-off between the consensus performance and the event-triggering frequency. As discussed in Remark 4, the condition \ud835\udefc2 > 0 is important for excluding the Zeno phenomena. For validating its effectiveness, let \ud835\udefc2 = 0; the event-triggered time instant of agent 1 is shown in Figure 13. Figure 13 shows that the event-trigger 1 is always violated after t > 8; it means that the unwanted Zeno phenomena appears. Because the threshold function becomes \ud835\udefc1e\u2212\ud835\udefdt, it converges down to zero quickly. Note that the function ||fi(t)|| is an increasing function; the event-triggering condition ||fi(t)|| < 0 cannot be guaranteed thus the event-trigger will be inevitably violated. That is the reason why the Zeno phenomena exhibits when \ud835\udefc2 = 0. Although the consensus of MASs can be achieved, however, the negative effective is that the Zeno phenomena appears when \ud835\udefc2 = 0." ] }, { "image_filename": "designv11_22_0000701_humanoids43949.2019.9035038-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000701_humanoids43949.2019.9035038-Figure3-1.png", "caption": "Fig. 3: Situation in which two non-coplanar surfaces of a link (shown in black) could make simultaneous contact with the environment (shown in blue).", "texts": [ " From this reasoning, the force control needs to be tackled in two steps: (a) We need to \u201cdrive\u201d the QP to produce on a given contact link a desired force and moment with respect to a given anchor point. In this way, a local CoP can be specified. (b) Then, we need to ensure the application of this wrench via force control. These objectives can be fulfilled by defining two additional tasks, which we named Wrench Task and Admittance Task. Let us consider a general case in which one or more non-coplanar surfaces of a link are in contact with the environment (see Fig. 3). It is then natural to associate the wrench task not to a surface (like in [10]) but to a link. We choose an anchor point on it, a, with respect to which we define the moment of the wrench. The relative position of the application of each lumped force with respect to this point, expressed in the global frame, is ri/a. Then, for k lumped forces acting on the link we have (see (2)):[ \u03b21 . . . \u03b2k S ( r1/a ) \u03b21 . . . S ( rk/a ) \u03b2k ] \u03c11 ... \u03c1k = [ fd na,d ] [ 0ini Wl 0fin ] \u03c1r = Fa,d. (8) Here, fd \u2208 R3 and na,d \u2208 R3 represent the desired resultant force and moment with respect to a of the link" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003079_0954407020947494-Figure7-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003079_0954407020947494-Figure7-1.png", "caption": "Figure 7. System responses of the controlled system (a2 = \u20133.000) when K1 = 100 and K3 = 50: (a) time history responses of the wheel hub speed and (b) phase diagram.", "texts": [], "surrounding_texts": [ "In order to verify the effectiveness of the designed torsional vibration stabilization controller for the wheelside direct-driven transmission system, the numerical simulation analysis is carried out in MATLAB 2015b environment. First, the ability of the cubic feedback control parameter on changing the system Hopf bifurcation type is verified. Figure 4(a) and (b) shows the time history responses of the wheel hub speed and the system phase diagram near the subcritical Hopf bifurcation point of the controlled system when K1=100 and K3=0, respectively. It is seen that when the bifurcation parameter is taken as a2=\u20133.100 which is close to the subcritical Hopf bifurcation point of the controlled system from the side of the stability region, the system responses are divergent without the cubic feedback control parameter introduced. Fortunately, when the cubic feedback control parameter is taken as K3=10, under the same initial values, the time history responses of the wheel hub speed and the system phase diagram near the subcritical Hopf bifurcation point of the controlled system are shown in Figure 5(a) and (b), respectively. It is obvious that the responses of the controlled system converge, and the bifurcation type of the controlled system at a2=\u20133.044 changes from the subcritical point to the supercritical point. It is shown that the introduction of the cubic feedback control parameter can effectively change the bifurcation type of the original system, and the theoretical analysis results are verified. Compared with Figure 5, Figure 6 shows the system responses near the subcritical Hopf bifurcation point of the controlled system with different initial conditions and same control gains. It is seen form Figure 6(a) that the system responses near the subcritical Hopf bifurcation point are also stable with the increase of the initial values, which is also found from the system phase diagram as shown in Figure 6(b). Thus, the abilities of the designed torsional vibration stabilization controller on adjusting the system stability region and changing the type of the system bifurcation point are proved. Furthermore, in order to clarify the influence of the cubic feedback control parameter of the designed torsional vibration stabilization controller on the system response amplitudes at the supercritical Hopf bifurcation point, the bifurcation parameter is taken as a2=\u20133.000 and the cubic feedback control parameters are taken as K3=50, 100, 200 for analysis. The system responses are shown in Figures 7\u20139, respectively, where Figures 7(a), 8(a), and 9(a) show the time history responses of the wheel hub speed and Figures 7(b), 8(b), and 9(b) show the system phase diagram. It is seen that within the effective range of the cubic feedback control parameter of the torsional vibration stabilization controller determined theoretically, with the increase of the cubic feedback control parameter, the responses amplitudes of the system supercritical Hopf bifurcation tend to decrease. Therefore, in practical application, larger cubic feedback control parameter of the torsional vibration stabilization controller should be adopted to eliminate the torsional vibration destabilization phenomenon of the wheel-side direct-driven transmission system and suppress the amplitudes of the stable responses." ] }, { "image_filename": "designv11_22_0001763_j.ifacol.2016.03.069-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001763_j.ifacol.2016.03.069-Figure1-1.png", "caption": "Fig. 1. Definition of forces, moment and angles", "texts": [], "surrounding_texts": [ "L = QSCL, D = QSCD, M = QSc\u0304CM (5) where, Q = 1 2\u03c1v 2 is the dynamic pressure, S is the projected wing area, and c\u0304 is the mean chord. CL, CD, and CM are the non-dimensional lift, drag, and pitching moment coefficients, respectively, which are approximated by the following equations for an aircraft with flaps. CL =CL0 + CL\u03b1\u03b1+ CL\u03b4e \u03b4e + CL\u03b4f \u03b4f (6) CD =CD0 + kC2 L (7) CM =CM0 + CM\u03b1 \u03b1+ CMq qc\u0304 2v + CM\u03b4e \u03b4e + CM\u03b4f \u03b4f (8) where, \u03b4e and \u03b4f are the elevator and flap deflection angle respectively. In case of aircraft with no flaps, the aerodynamic model for Lift and Pitching Moment do not contain \u03b4f term. For thrust (FT ) we use the following model, which is valid for propeller driven electric aircrafts. FT = 1 2 Th(\u03c1SpropCprop + v2/kmotor 2) (9) where, Th is the throttle input, Sprop and Cprop are propeller parameters, and kmotor is motor parameter. The objective is to design a longitudinal control to follow the references velocity and altitude commands precisely under the parameters uncertainty of the UAV model. 3. PRELIMINARY In this section, we present general theory of nonlinear dynamic inversion(NDI)(Enns et al. (1994)) and neuroadaptive approach (Padhi et al. (2007)) in brief. The approaches will be use in the next section to develop an adaptive longitudinal controller of UAVs." ] }, { "image_filename": "designv11_22_0002801_s11661-020-06013-7-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002801_s11661-020-06013-7-Figure1-1.png", "caption": "Fig. 1\u2014(a) A schematic of the self-designed cuboctahedron unit cell; (b) The three-layered architecture of the specimen for the uniaxial compression testing used in this study.", "texts": [ " The second objective of this study was to examine the effects of HIP treatment on the compression performance and fracture mode. In-situ mechanical testing coupled with digital image correlation (DIC) technique was used to investigate in detail the progress of deformation and fracture under uniaxial compressive stress. The correlations between microstructure, compression properties, deformation behavior, and fracture mechanisms are discussed. The self-designed cuboctahedron unit cell proposed in a previous study[27] is shown in Figure 1(a). In the structure, three straight struts lie along the X, Y, and Z axes, and four struts are inclined at 45 deg. The dimensions in the X, Y, and Z axes of this cell are 2, 2, and 1.44 mm, respectively. This new cuboctahedron unit cell was designated as a COH-Z structure.[31] For the uniaxial compression test, cylindrical specimens with a sandwiched architecture were fabricated, as shown in Figure 1(b). The diameter of the cylindrical samples was 13 mm. The sandwiched cylinder contained a 5 mm-thick top solid section, a 15 mm-thick medium cellular section, and a 5 mm-thick bottom solid section. The total height of this cylinder was 25 mm. In the original design, the strut diameter in Figure 1 was 250 lm. After SLM, the average strut diameter was enlarged to 310 lm. To prepare the sandwiched samples, a plasma-atomized Ti-6Al-4V spherical powder with a median size (D50) of 34 lm was used.[9] The specimens in Figure 1(b) were prepared in argon with SLM equipment (SLM 250 HL, SLM Solutions GmbH, Lu\u0308beck, Germany). A laser power of 225 W, a scan velocity of 658 mm/s, a layer thickness of 50 lm, and a hatching spacing of 120 lm were applied to produce the sandwiched samples in this study. In each successive layer, the scanning direction was rotated by 75 deg. The above process parameters were recommended by SLM Solutions GmbH. To identify the influences of microstructure on compression performance, some SLM specimens were HIPed in a hot isostatic press manufactured by Avure Technologies" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002448_j.cirp.2020.04.060-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002448_j.cirp.2020.04.060-Figure3-1.png", "caption": "Fig. 3. (a) Schematics of the baseplate and foam stainless block, (b) multi-layer type deposition, (c) cross-layer type deposition.", "texts": [ " Considering that the pores disappear in the higher layer, the bubble easily collapses due to overheating. In addition, our previous work showed that the porosity rate is low in a onelayer deposition because the heating time is too short to generate pores. In order to produce a foam stainless block with high porosity, it is essential to consider the deposition track and laser power control, to avoid overheating and to ensure a sufficient amount of melting time. Therefore, this study proposes a special nozzle trajectory for foammetal DED, as shown in Fig. 3. To generate a sufficient number of pores inside the deposit, several layers should be deposited continuously; thus, a 4- or 6-layer line deposition is conducted in parallel with a 4 mm pitch in order in this test. Furthermore, the deposition direction is also changed in the series of upper layers in the crosslayer type deposition, which can help suppress the anisotropic characteristics. Furthermore, a closed-loop control (CLC) is also employed Table 1 Content ratio of SUS316L powder (excepting TiH2 and Te)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000349_ilt-03-2019-0088-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000349_ilt-03-2019-0088-Figure2-1.png", "caption": "Figure 2 Schematic diagram of boundary slip", "texts": [ "2Modified Reynolds equation The film thickness is small relative to other dimensions of length, to simplify the followed numerical analysis, basic assumptions for themodel aremade as follows (Lebeck, 1991): the lubricant is Newtonian and the flow is laminar; the film composed of a biphasic mixture is divided into full liquid film zones and cavitation zones and the pressure in the cavitation zones remains constant; thermal wedge of the fluid and thermal distortion of the friction pair are neglected; the angular misalignment of the friction pair and the fluid inertial are neglected; and the model is suitable for the isothermal and adiabatic conditions. Based on the assumptions, the Navier\u2013Stokes equations are simplified as follows: @p @x \u00bc @ @z m @u @z @p @y \u00bc @ @z m @v @z @p @z \u00bc 0 8>>>>>< >>>>>: (1) The Navier slip model is shown schematically in Figure 2. The velocity of the liquid at the wall-related to the slip length can be expressed as follows (Navier, 1823): uslip \u00bc b @u @z wall (2) where b is slip length, mm, b is defined as the ratio of the slip velocity to the absolute value of the velocity gradient in the normal direction of the wall; uslip is the fluid velocity on the wall. The choice of Navier slip model is not only for its Improved hydrodynamic performance Yun-Lei Wang, Jiu-Hui Wu, Mu-Ming Hao and Lu-Shuai Xu Industrial Lubrication and Tribology Volume 71 \u00b7 Number 9 \u00b7 2019 \u00b7 1108\u20131115 simplicity but also it is the most common slip model that shows good correlation to experimental results (Neto et al" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.45-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001889_b978-0-08-100036-6.00002-9-Figure2.45-1.png", "caption": "FIGURE 2.45 Schematic outline of the brush-string model.", "texts": [ " Two deflections are distinguished in longitudinal and lateral directions, which we shall denote as u and v, respectively. These deflections refer to the tread deflection and the belt deflection, indicated with subscripts t and b, respectively: u5 ut 1 ub; v5 vt 1 vb \u00f02:84\u00de Only the variation of deflections with the longitudinal coordinate x is studied here, i.e., all deflections and forces are assumed the averaged value over the width of the contact area. Conditions such as turn-slip of very large camber angles, for which this assumption is not correct, are not considered. A schematic outline is shown in Figure 2.45. The belt is modeled as an infinite string under a tension force and is connected to the tire symmetry plane through longitudinal ccx and lateral carcass stiffnesses ccy per unit length. In the contact area, brush elements are attached to the belt. It is assumed that both longitudinal and lateral slip are present, with resulting shear forces Fx and Fy and aligning torque Mz. The leading edge of the adhesion part of the contact area (usually equal to the leading edge of the entire contact area) is displaced with belt deflections u1 and v1 in longitudinal and lateral directions, respectively. Because of the presence of carcass stiffnesses per unit length ccx and ccy, any part of the belt with length dx experiences resistance forces in x and y direction, equal to ccx u dx and ccy v dx, respectively. With the definition of qx and qy according to Eq. (2.80) (i.e., being the longitudinal and lateral external forces on the tire per unit length) and considering the equilibrium of a part of the belt with length dx (see Figure 2.45), one is able to derive the following differential equations for the belt defections ub and vb S1 d2ub dx2 2 ccx ub 52qx; 2a, x, a \u00f02:85a\u00de S2 d2vb dx2 2 ccy vb 52qy; 2a, x, a \u00f02:85b\u00de where S1 is the longitudinal elastic resistance of the tread band (Young\u2019s modulus E times cross sectional area) and S2 is the effective tension force in the belt. We refer to Pacejka [31] and [32] and Higuchi [17] for more details about the derivation of Eqs. (2.85a) and (2.85b). Clearly, the Eqs. (2.85a) and (2.85b) are not restricted to the contact area, but hold on the entire belt, where the noncontact part is \u201crolled out to infinity.\u201d Considering Eq. (2.85b) for qy5 0, and assuming the belt deflection to be finite for x. a, one finds from integration for x. a that vb\u00f0x\u00de5C e2x ffiffiffiffiffiffiffiffiffi ccy=S2 p ; x. a and therefore, at x5 a dvb dx \u00f0x5 a\u00de52v1 ffiffiffiffiffiffi ccy S2 r 52vb\u00f0x5 a\u00de ffiffiffiffiffiffi ccy S2 r \u00f02:86\u00de Consider Figure 2.45, in which we indicate the distance \u03c3\u03b1 between the leading edge of the contact area and the intersection of the line through the straight part of the contact zone (adhesion) with the wheel center plane. This distance is referred to as the relaxation length for the belt deflection in the lateral direction. Further, it is related to the distance a tire needs to travel before a significant percentage of steady-state shear force is reached following a sudden change in the slip angle. With the line through the straight part of the contact zone tangent to the deflection profile, one should have dvb dx \u00f0x5 a\u00de52 v1 \u03c3\u03b1 \u00f02:87\u00de A similar reasoning holds for longitudinal deflection, in which we introduce the relaxation length \u03c3\u03ba for the belt deflection" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000956_ijmmme.2015100103-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000956_ijmmme.2015100103-Figure2-1.png", "caption": "Figure 2. Calorific groove", "texts": [ " The gradients in the groove of the bowl are explained same manner. At the beginning of braking, the temperature of the bowl is 20 \u00b0C while that of the tracks is a few hundred degrees. Moreover, in order to prevent the temperature of the hub is not too high (what would generate Copyright \u00a9 2015, IGI Global. Copying or distributing in print or electronic forms without written permission of IGI Global is prohibited. rises in temperature of the tire, very critical for its behavior), the groove is machined so as not to transmit too much heat to the bowl (Fig.2). With this machining, the temperature of the bowl effectively decreases, but the heat gradients increase consequently in this zone. Those generate thermal stresses which explain the ruptures of bowl observed during severe experimental tests. There are two types of disc: full discs and ventilated discs. The full discs, of simple geometry and thus of simple manufacture, are generally placed on the rear axle of the car. They are composed quite simply of a full crown connected to a \u201cbowl\u201d which is fixed to the hub of the car (Fig" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003327_s0263574720001484-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003327_s0263574720001484-Figure4-1.png", "caption": "Fig. 4. Spatial configuration of an inter-module.", "texts": [ " 3 can be considered as a rigid parallel robot composed of k independent kinematic chains whose ends are connected by two rigid https://doi.org/10.1017/S0263574720001484 Downloaded from https://www.cambridge.org/core. University of Toledo, on 27 May 2021 at 11:28:14, subject to the Cambridge Core terms of use, available at https://www.cambridge.org/core/terms. platforms, namely the lower and upper modules. The lower ends of the joints linked to the lower module form a regular polygon, as shown in Fig. 4. This is also the case for the upper ends linked to the upper module. These joints have variable lengths and are represented by li, j . They provide the position and orientation of the upper module relative to the lower module. The variable represented by d j is the distance between the centers of gravity of the lower and upper modules; it is considered as a passive joint. The orientation represented by \u03c8 j , \u03b8 j and \u03c6 j are the roll, pitch, and yaw angles, respectively. The inter-module is then modeled as a parallel robot with k universal-prismatic spheric and one universal-prismatic (kUPS-1UP joints)" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003550_j.matdes.2021.109793-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003550_j.matdes.2021.109793-Figure1-1.png", "caption": "Fig. 1. Tantalum-titanium biomedical screw implant cross-section concept rendered in CAD showing the critical interface between two separate materials and structures.", "texts": [ " Our results indicate that the underlying temperatures achieved in bimetallic components can be predicted accurately using a detailed FEA approach, aiding larger-scale approaches towards improved process modeling and manufacturing of bimetallic structures using laser-based AM. 2021 The Author(s). Published by Elsevier Ltd. This is an open access article under the CC BY-NC-ND license (http://creativecommons.org/licenses/by-nc-nd/4.0/). Additively manufactured (AM) bimetallic structures, such as the Ti-Ta screw-in Fig. 1 that consist of a single component with a defined transition from one material to another, are highly desirable in many industries due to combining the best of multiple metals in a single structure [1\u20133]. While exciting from an application standpoint, these structures are extremely challenging to process via traditional manufacturing methods like welding, tape-casting, diffusion bonding, brazing, etc., and motivate the investigation of alternative processing routes such as AM. An example of a potential application is shown in Fig. 1, where a biomedical bone screw is composed mostly of titanium due to corrosion resistance, fatigue resistance, and high strength-weight ratio relative to most other materials [4], with a structurally graded tantalum outer surface to promote increased bioactivity in comparison to titanium [5]. Additionally, tantalum is a refractory material (melting temperature > 3000 C) used extensively in the chemical processing industry due to its enhanced corrosion resistance in alkaline and mineral reducing acids [4,6] and has shown some promise being processed via laser-based AM [7\u20139]" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000035_978-3-319-96968-8_13-Figure13.5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000035_978-3-319-96968-8_13-Figure13.5-1.png", "caption": "Fig. 13.5 a Defocusing in up, b focusing position, and c defocusing in bottom", "texts": [ "86 Characterization of laser marking may be modified for change in the marking parameters; therefore, choosing of the proper laser marking parameter and controlling of the laser marking parameters are required. This paper deals with laser process parameters modeling like power setting, duty factor, pulse frequency, speed defocus height on the mark characteristics. Relatively uncommon process variable, i.e., defocus height, acts a very major role in determining the mark characteristics. It is clearly explained below in Fig. 13.5a\u2013c about defocusing in up, focusing position, and defocusing in bottom, respectively. Figure 13.5 shows that kerf width is more in defocus up rather than in focus point. The kerf width increases with the increase in defocus height up, and with respect to power, it increases more steeply at higher power. It is also observed that, for the same parameter setting, in defocus up (\u2212), the kerf width is more than defocus in bottom (+). At defocus bottom, where kerf width increases slightly or almost same for higher power also. Laser beam machining is all about thermal energy transformation, which causes melting and direct vaporization of material" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000997_tmag.2015.2443026-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000997_tmag.2015.2443026-Figure2-1.png", "caption": "Fig. 2. (a) Schematic of a three-axis electromagnet fitted on an optical microscope. (b) Schematic of disk-shaped particle rotation in a Newtonian fluid under a rotating magnetic field.", "texts": [ " The suspension 0018-9464 \u00a9 2015 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission. See http://www.ieee.org/publications_standards/publications/rights/index.html for more information. has a very low volume percent of 0.001 to minimize interactions between the microdisks. The motion of an isolated microdisk is observed under an optical microscope fitted with a three-axis electromagnet. The angle between the microdisk plane and in-plane rotating field is defined as \u03b8 [Fig. 2(b)]. The microdisk is first aligned perpendicular (\u03b8 = 90\u00b0) to the image plane by a constant magnetic field. Subsequently, an in-plane rotating magnetic field is turned on to align the microdisk parallel (\u03b8 = 0\u00b0) to the image plane. The microdisk rotation is captured by a high-resolution microscope camera (1292 \u00d7 964 resolution, 30 frames/s) under different aligning field strengths, field rotation frequencies, and fluid viscosities. Five different microdisks are imaged for each experimental condition to confirm the repeatability of measurements", " The formation of microdisk layers is attributed to the magnetic dipole attraction between neighboring Ni-Fe microdisks. Such chaining is undesirable as it reduces the electrical resistance and increases eddy current losses in the material, motivating the present effort to investigate the planar alignment process by systematically varying the rotating field strength, field rotation frequency, and matrix viscosity. The rotation behavior of the disk-shaped particles in a fluid of known viscosity is observed under an optical microscope using a 40\u00d7 objective lens [Fig. 2(a)]. The microscope is fitted with a three-axis electromagnet for applying fields while imaging. A theoretical model is developed to describe the hydrodynamic behavior of disk-shaped particles in a Newtonian fluid at small Stokes numbers. As shown in Fig. 2(b), there are three torques acting on an isolated disk-shaped particle: 1) magnetic torque from the rotating field, Tm ; 2) drag torque from the fluid viscosity, Td ; and 3) Brownian motion torque, Tb. The magnetic torque, Tm , orients the easy plane of the disk-shaped particle parallel to the field rotation plane; while the fluid drag torque, Td , and Brownian motion torque, Tb, act against this motion. The microdisk is approximated as an oblate ellipsoidal particle to simplify fluid drag torque calculation" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002523_j.matpr.2020.05.322-Figure19-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002523_j.matpr.2020.05.322-Figure19-1.png", "caption": "Fig. 19. CATIA Model of Tubular Anti-Roll Bar.", "texts": [ " 14 to 16 shows X, Y, Z axis strain displacement. Another strategy is to list the outcomes \u2018\u2018DOF Solution - UY dislodging\u201d. The acquired uprooting esteem is put away for use in move firmness count that will be introduced in the accompanying segment. Presently, the consequences of the subsequent burden venture of the static investigation can be perused to the database. Figs. 17 and 18 Equivalent Strain and Von mises strain of Solid Bar. Here, the anxiety circulations on the bar under most extreme stacking are the contemplations. Fig. 19. shows CATIA model of tubular anti-roll bar. This outline shows the initial five regular frequencies of the counter move bar. Not the same as the static investigation, here assurance of every common recurrence is viewed as an alternate burden step. Figs. 20 to 27 shows the meshed model view and the carbon/epoxy resin anti roll bar. Consequently, the most extreme qualities at each segment of the bar can be resolved out- the load Constrained Model. Fig. 21. Von-Mises Stress Analysis. Fig. 22. Carbon/Epoxy Resin Anti roll Bar \u2013 Stress Plot in Z \u2013 Axis" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000699_r10-htc47129.2019.9042460-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000699_r10-htc47129.2019.9042460-Figure3-1.png", "caption": "Fig. 3 3D Model of Stove Design", "texts": [], "surrounding_texts": [ "In this section, we will present the specific points that were covered using the above mentioned co-design tools that lead us to the final design. We have also given a description of the design we propose for the cook stove." ] }, { "image_filename": "designv11_22_0002452_rspa.2020.0062-Figure9-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002452_rspa.2020.0062-Figure9-1.png", "caption": "Figure 9. Geometry of the typical segment of the tensegrity beam: (a) linear case; (b) sub-linear case; (c) super-linear case. The Autocad files used to 3D print the beam prototypes are linear.dwg, sublinear.dwg and superlinear.dwg for the linear, sublinear and super-linear cases, respectively. These have been provided in the electronic supplementary material. (Online version in colour.)", "texts": [ " The average difference with respect to the theoretical value is always less than 0.6% within the range 0 \u2264 \u03d5i \u2264 0.5 rad, in particular, less than 0.01% for the linear case. (b) Manufacturing of prototypes The prototypes were manufactured with a Prusa 3D printer in polylactic acid, a thermosoftening plastic obtained from renewable resources such as corn, using the deposition method (fused deposition modelling) [13]. The typical segments for the three considered cases, with an indication of their size, are represented in figure 9. Each beam, consisting of 10 segments, has a free span of 613.5 mm, calculated from the midpoints of the left- and right-hand side supports in the reference undeformed configuration, with a total weight of 2.313 N, 2.948 N and 2.403 N for the linear, sub-linear and super-linear cases, respectively. The prestressing cable is a polyamide 6.6 wire of diameter 1.1 mm, placed in series with one spring of stiffness k0 = 3.9 N mm\u22121 at each end, as represented in figure 10. In this way, the effective stiffness K of the system, introduced in (2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003327_s0263574720001484-Figure8-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003327_s0263574720001484-Figure8-1.png", "caption": "Fig. 8. Compact Bionic Handling Assistant(CBHA).", "texts": [ " and \u23a7\u23aa\u23a8 \u23aa\u23a9 f \u2212 sign( f ) f = atan2(r f 23, r f 13) or f \u2212 sign( f ) f = \u2212atan2(r f 23, r f 13) if f = \u00b1\u03c0 2 In this section, we focus on the validation of the proposed FKM scheme. It begins with a description of the experimental platform, followed by the database generation process and the results of the training of each architecture. The section concludes with a comparison with the FKM models previously developed, namely model-based and learning-based approaches and the MLP-hybrid approach. The proposed scheme is validated using a Compact Bionic Handling Assistant (CBHA) continuum manipulator, as shown in Fig. 8. It is a compact version of the Bionic Handling Assistant (BHA) robot16 and is inspired by the elephant\u2019s trunk. CBHA is both economical and lightweight, thanks to its additive manufacturing process in which printing is carried out by polyamides. It consists of two bending sections, each equipped with three pneumatic actuators, a wrist axis, and a compliant gripper. A PID regulator controls the pressure supply in each tube. The elongations of the different tubes are provided by six-wire cable potentiometers placed along each tube" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001533_0142331215619972-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001533_0142331215619972-Figure2-1.png", "caption": "Figure 2. Definition of _xR during a steady turn.", "texts": [ " Let WP1 and WP2 be two consecutive waypoints on the earth\u2019s surface on the desired path, and let x R be the angle of the line WP1-WP2 with respect to north (called the reference or desired course angle). This is computed using latitude and longitude information of the waypoints (Samar et al., 2013). Its rate of change, i.e. _x R is zero for straight paths, and can be calculated for circular paths as: _x R = V R . The course angle x is the angle of the ground velocity V with respect to north. Another important parameter is x E =x x R , called the intercept course. For circular path following (Figure 2), consider WP1 and WP2 as two consecutive points on an arc with center of turn O and radius of turn R. Point P is the nearest point on the arc to the actual position of the vehicle. The reference course x R is the angle of the tangent line at point P with respect to north. The main task of the guidance algorithm is to keep the cross-track error y as small as possible, and also to keep x E \u2019 0 when y \u2019 0. In case of a non-zero y, the guidance algorithm will manipulate x E by banking the vehicle to bring y to zero" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0003578_iccmc51019.2021.9418287-Figure4-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0003578_iccmc51019.2021.9418287-Figure4-1.png", "caption": "Figure 4. Ultrasonic Sensor.", "texts": [], "surrounding_texts": [ "This module is the combination of both GSM and GPS system. It basically works on the AT commands. To determine the exact location of the vehicle or boat on the earth can be done using GPS, here it will send the signals to satellites and to the earth ground stations, and receives information like A ltitude, Longitude, Lat itude [4], Time, etc. in the form of NMEA string. By using this informat ion, the exact location of the boat can be identified, the SIM808 GPS receiver module uses USART communicat ion to communicate with Arduino or PC Terminal. Global System for Mobile Communicat ion (GSM) [9] is uses that extracted in formation of location from GPS is sent to the respected mobile numbers by processing the code present in the Arduino by following the AT commands in code by GPRS System." ] }, { "image_filename": "designv11_22_0001311_iccas.2014.6988027-Figure1-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001311_iccas.2014.6988027-Figure1-1.png", "caption": "Fig. 1 Physical model of a quadrotor", "texts": [ " In Sec tion 2, we describe the dynamic model of quadrotor UAV used in this study. Section 3 describes the design of pro posed adaptive control strategy and altitude compensa tion scheme. Simulation results are presented in Section 4, and Section 5 draws concluding remarks and presents future work. In this section, we examine the dynamics of the quadrotor helicopter in order to gain the insight neces sary for the controller design. Quadrotors consist of four rotors attached to a rigid cross airframe as shown in Figure 1, with two opposing rotors rotating clockwise (1,3) and the other two rotating counterclockwise (2,4). Control of quadrotor is achieved by differential control of the thrust generated by each ro tor. Vertical motion is accomplished by simultaneously increasing or decreasing the speed of all four rotors. Pitch motion is achieved by difference in speed of the front rear set of rotors. Roll motion is archived in the same way using left-right set of rotors. Quadrotor is an un deractuated system, which means that forward/backward and left/right motions are coupled with pitch/roll motions respectively and can be controlled through them" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000514_j.mechmachtheory.2019.103667-Figure20-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000514_j.mechmachtheory.2019.103667-Figure20-1.png", "caption": "Fig. 20. Comparison of compensated geometry of face gear and standard geometry generated by a shaper with 26 teeth.", "texts": [ " In this section, alternative methods to prepare the face-gear tooth surfaces to absorb the errors of alignment due to shaft deflections are presented. Two methods have been investigated. The first one is based on application of an offset between the axis of rotation of the shaper and the face gear, which is equivalent to the compensation of E, the minimum distance between shafts. The second one is based on the application of a shaper with 28 teeth to increase the longitudinal crowning applied to the face-gear tooth surface and therefore to make the face-gear set more insensitive to errors of alignment. Fig. 20 shows the results of comparison of the compensated geometry of the face-gear tooth surface after the third iteration of compensation (see Table 3) with respect to the standard geometry achieved with the generation of the face gear by a shaper with 26 teeth and no errors of alignment compensated. Notice that the deviations obtained after compensation of errors of alignment caused by shaft deflections show a longitudinal deviation pattern with a maximum of 140 \u03bcm at the inner radius of the face-gear tooth surface", " If the face width of the face gear is denoted by Fw2 and the required deviation of the face-gear tooth surface is \u03b4max at the inner radius (point A2 in Fig. 21), the shaft of the shaper has to be rotated around point A1 of an angle \u03bb given by \u03bb = arctan \u03b4max F (29) w2 Point A1 is located at the outer radius of the face gear denoted by L2. Therefore, the applied error E, as shown in Fig. 21, is easily determined as E = L2 tan \u03bb = L2 \u03b4max Fw2 (30) Considering the data shown in Table 1 for the face gear and a desired deviation at the inner radius of the face gear of 140 \u03bcm as shown in Fig. 20, the error E to be compensated is of 0.89 mm. For the convenience of identifying different designs, the reference design with all errors compensated, that is to say the face gear with compensated geometry for E = 1.2923 mm, A1 = \u22120.0188 mm, A2 = 0.4917 mm, and \u03b3 = \u22120.0304\u25e6, will be referred to as Case 0. The design in which only an error E = 0.89 mm is compensated will be referred to as Case 1. Later, the case of design in which the face gear is generated by a shaper with 28 teeth and without errors of alignment compensated during generation will be referred to as Case 2" ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0002306_j.matpr.2019.08.034-Figure2-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0002306_j.matpr.2019.08.034-Figure2-1.png", "caption": "Fig. 2 Fabrication of PMGs (a) 3D-printed PMGs using Meta Bolt V2 3D-printer and (b) Configuration setting on Repetier Software.", "texts": [ " The rotor is composed of windings that are connected to energy storage devices. In this case the rotor has 12 slots. The parameters of designed PMGs are concluded in Table 1. For the generator fabrication, we use a Meta Bolt V2 3D-printer as shown in Fig (a). In this work, the most common 3D printing material, Polylactic Acid (PLA) is used as 3D printing filament. The reason to use PLA filament because it is easy to use and is made from renewable resources and thus, biodegradable. The printer is controlled via Repetier Software as presented in Fig 2 (b). The temperature of the nozzle and heating bed are set to 220 C\u030a and 60 \u030aC, respectively. Normal PLA filament with 1.75 mm of diameter is used to make stator and rotor. The layer height or layer resolution and printing speed are chosen to be 0.2 \u00b5m and 50 mm/s respectively. Moreover, thin supports with a height of 0.2 mm are included via 3D printing in order to mount the generator. These supports are thin enough to have a negligible effect on the generator. The construction of the PMGs begins with constructing the coil winding " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0000574_978-981-15-1263-6_7-Figure6.10-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0000574_978-981-15-1263-6_7-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_22_0001760_j.mechmachtheory.2016.03.023-Figure3-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001760_j.mechmachtheory.2016.03.023-Figure3-1.png", "caption": "Fig. 3. 4R mechanism, joint axes a\u03021, . . . , a\u03024.", "texts": [ " , n \u2212 1, and the n \u2212 1 linear factors k\u2032 l, l = 1, . . . , n \u2212 1. They are summarized to x = {a\u03021, . . . , a\u0302n\u22122,k,k\u2032}. Compared to Eq. (26), the number of equations as well as the number of unknowns are reduced by 14 \u2212 2n. Example: 4R mechanism(BENNETT). In the following joint axes of a 4R loop are determined that fulfil the 1st- and 2nd-order closure conditions. Without loss of generality the joint axes a\u03021 = [ 1 0 0 0 0 0 ]T , a\u03022 = [ a21 0 a23 a23 0 \u2212a21 ]T , \u221a a2 21 + a2 23 = 1 (34) are defined according to Fig. 3 immediately fulfilling the normalisation and PL\u00dcCKER conditions, refer to Eq. (69) in the Appendix. From the solution of the second-order closure condition in Eq. (32) with n = 4 the third joint axis a\u03023(k,k\u2032) = ( \u02dc\u0302a1 k1 k3 + \u02dc\u0302a2 k2 k3 \u2212 E k\u2032 3 \u22121 ( a\u0302\u2032 2 k2 + a\u03021 k\u2032 1 + a\u03022 k\u2032 2 ) (35) is expressed in terms of k,k\u2032. Remark: There exists one special position for the BENNETT mechanism, where det(\u0303a\u03021 k1 k3 +\u02dc\u0302a2 k2 k3 \u2212 E k\u2032 3) = 0 holds. In this position the joint axes fulfil the special condition in Eq", " By this the fulfilment of the closure conditions up to the maximum order m = 2 is a sufficient condition for the mobility of the BENNETT mechanism. This condition also holds for the mobility of the planar and the spherical 4R mechanism. Outgoing from the given axes a\u03021, a\u03022 from Eq. (34) with the distance = 0 between the first and the second joint axes and arbitrary joint angles b12 = 0 and calculating axes a\u03023, a\u03024 with Eq. (35), Eq. (36) together with the conditions Eq. (38) lead to a spherical 4R mechanism. Accordingly angle b12 = 0 between a\u03021, a\u03022 corresponding to a23 = 0 in Fig. 3 and arbitrary values of the distance = 0 lead to a planar 4R mechansim. Obviously the spherical and the planar 4R mechanism have finite mobility, see also the last example in Subsection 4.2. Thus, the fulfilment of the 1st- and 2nd-order closure conditions is sufficient for the finite mobility of 4R mechanisms, including the spherical, the planar, and the BENNETT 4R mechanism. 4.2. Special analytical solution for every order m In the following special solutions of the mobility conditions in Eq. (26) are presented which guarantee that all higher order mobility conditions are fulfilled, gm(a\u0302i) = 0, m = 1, " ], "surrounding_texts": [] }, { "image_filename": "designv11_22_0001513_j.proeng.2015.12.642-Figure5-1.png", "original_path": "designv11-22/openalex_figure/designv11_22_0001513_j.proeng.2015.12.642-Figure5-1.png", "caption": "Fig. 5 Anisotropic FEA for mini-ruler under tension loading and different nominal fiber orientations: a) FO=0\u00b0 b) FO=30\u00b0 c) FO=90\u00b0", "texts": [ " A mean aspect ratio of about 27 was chosen. Finally an orthotropic material model is created and used by the finite element software Abaqus\u00ae [8]. Tab. 1 Material properties [2] Material Young\u2019s modulus E/ MPa Poisson\u2019s ratio \u03bd Density \u03c1/ (g/cm\u00b3) Matrix (PA66) 3100 0.40 1.16 Glass fibers 72000 0.22 2.54 The fiber orientation distribution over the thickness (skin-core effect) affects the stress distribution. This effect is exemplarily shown for the Mini-Ruler-Specimens with different nominal fiber orientations in Fig. 5, where the maximum principle stress is printed as a contour plot. The highest stress value occur at the surface layer for longitudinal and at the core layer for transversal fiber orientation. At FO=30\u00b0, for the stress distribution a sharp peak value at the center of the core layer is observed. The stress concentration factors Kt based on the maximum principle stress for the anisotropic FEA are presented in Tab. 2. Remarkably, in contrast to the isotropic case, the stress concentration factor for torsion loading is higher than for tension loading" ], "surrounding_texts": [] } ]